1 0 JAN. 1974

RCHIEFDOKTORSAVHANDLINGAR Lab. v. Scheepsboiniikunt

Technische Hogeschigit,

Deift

VD

CHALMERS TEKNISKA. HOGSKOLANr 11 9

.OHALMERS.-- UNIVERSITY OF TECHNOLOGY:GOTEBORG.-'SWEDEN'.

A STUDY OFCRASH STOP TESTS WITH

SINGLE SCREW SHIPS

RALPH A. NOR RBY

INSTITUTIONEN FOR SKEPPSHYDROMEKANIK1972

C74613,...

A STUDY OF

CRASH STOP TESTS WITH

SINGLE SCREW SHIPS

RALPH A NORM

1972

A STUDY OF

CRASH STOP TESTS WITH

SINGLF SCREW SHIPS

BY

RALPH A NORRBY

AKADEMISK AVHANDLING

SOM MED TILLSTAND AV CHALMERS 'TEKNISKA HOGSKOLA

FRAMLAGGES TILL OFFENTLIG GRANSKNING FOR

TEKNOLOGIE DOKTORSGRADS VINNANDE FREDAGEN DEN.

26 JANUARI 1973, KL 10.00 A PALMSTEDTSSAL-FN VID

CHALMERS TEKNISKA HOGSKOLA, SVEN HULTINS GATA,

-GOTEBORG

SYNOPSIS

Over the next ten year period the volume of transport by

ship is expected to be doubled. The number of vessels will

increase which will accentuate the already difficult

situation in congested waterways. The medium and maximum

size of tankers will tend to increase as will the speed of

container ships. There is a definite requirement to

improve the manoeuvring qualities and also to simplify the

data available for appraisal of same on board. The paPer

deals with the stopping qualities of ships.

Means and methods to improve the stopping qualities are

discussed and devices such as brake flaps and controllable

pitch propellers as well as the method of rudder cycling

are dealt with. It is expected that modern ships will to

an increasing extent be equipped with devices to improve

the stopping qualities.

The data available on board usually consists of recordings

of stopping results at one or two displacements. This data

is difficult to interpret for intermediate conditions. The

Oosterveld diagrams are discussed as those are considered

rather practical. However, they only give the master

information on the ship's expected mean stopping qualities.

A large number of methods to calculate the stopping data

for ships both with fixed pitch propellers and controllable

pitch propellers have been suggested. Complicated conditions

prevail during a stopping manoeuvre and a number of simpli-

fications have therefore been introduced in the calculating

methods. In spite of these simplifications it is claimed

by several authors that the results from actual crash stop

manoeuvres conform well with calculated results.

1

In this work crash stop manoeuvring has been treated on the

basis of data from full scale tests. A relatively large

2

number of tests from ships with fixed pitch as well as

controllable pitch propellers form the basis of the

investigation, covering the range from coastal vessels

to large tankers. The data has been received from ship-

yards in various countries. Simple statistical methods

have been used to. deal with the material.

Observations from crash stop tests suggest that the _vessel's

course stability during the manoeuvre is of importance.

A. course stability criterion is developed for trimmed and

untrimmed vessels during crash stops. The influence,,of

propeller type, wind and direction of propeller rotation

on the sheer of a ship during this manoeuvre is discussed

as is also_the probability of ships sheering against each

other when making the crash.st0P manoeuvre. The turning

-angle for-ships with .fixed pitch propellers and with

controllable pitch propellers is investigated..

To determine track reach and stopping time simple relations

are given for the retardation factor. The influence of

retardation from ship hull resistance, turning, wind,

propeller and uncertain influences is discussed. Non-

dimensional numbers for the stopping distance and time

have been evaluated.

Propeller characteristics at

results based on atmospheric

are collated. Tests from the

introduced for evaluation of

of ships._

crash stop from published

tests at bollard pull astern

KMW cavitation tunnel are

the stopping characteristics

In evaluating. the stopping distance and time the statistical

material is treated with linear regression analysis. As

a result diagrams for estimating and appraising crash stOp

performance have been obtained for ships with fixed pitchand controllable pitch propellers.

In order to examine the proposed method to calculate the,:

stopping distance and time a comparison has been made

using a number of wellknown methods. For these the same

statistical material has been used throughout. It will be

seen that the proposed method, based on statistical

analysis, is at least as accurate as the best of the

other methods used for the stopping material investigated.

Stopping equations are developed which state the probability

that the stopping performance is inferior to the calculated

values. Head reach as well as lateral reach are discussed.

The use of practical stopping diagrams is demonstrated.

Several authors have made comparisons of stopping perfor-

mance of ships with fixed pitch propellers and with control],-1-

able pitch propellers. This data is compared with the

results based on the statistical material.

SYMBOLS

Symbol Definition Physical dimension

A frontal area above water

AEblade area

A constant, es' (41)

Atconstant, eq (42)'

A0propeller disc area

a constant, eq (44),

at constant, eq (45)

BM'moulded beam

Bs constant, eq (41)

Bt.conStant eq (42)

bs constant, eq (44)

btconstant, eq'(45)

4

Symbol Definition Physical dimension

centrifugal force,-2

longitudinal component

astern thrust constant,eq (33)

block coefficient

CQPtorque coefficient, eq (32) -

C thrust coefficient, eq (31) -TP

C constant, eq (3) -Y.v

C1constant, eq (4)

C2 constant, eq (4)

c1constant, eq (34)

c wind resistance coefficient, -X longitudinal

@ Froude resistance constant -

CPP controllable pitch propeller -

D propeller diameter

DE diesel engine

,d mean draught L

dAaft draught L

dFforward draught L

F total braking force MLT-2

FPP fixed pitch propeller -

acceleration of gravity LT2-,

,

h water depth L ,

I immersion of propeller shaft L

KQtorque coefficient *wn2D5

KTthrust coefficient Thivn2D4

MIM

kxquotient between longitudinally"added mass" and the vessel'smass

quotient between athwartships"added mass" and the vessel'smass

Lpp length between perpendiculars

-5-

Symbol Definition Physical dimension

N total number ofobservations

-1Nr moment-yaw velocity MI2 T

derivative

N' Nr/0,5pwlapp5d

Nvmoment-sway velocity MITderivative

N;N' Nv/0,5pwV1Ipp2 d

constant, eq (20) -

number of observations -

T-1propeller shaft sped-

11Areversed propeller. shaft T1speed

propeller shaft speed,0 design

tOshaft speed at stop order. Tn

propeller pitch L

PAastern propeller pitch at L

0,7 radius

17 statistical m.v.of PA LA

propeller design pitch at L0 0,7 radius

-1 --

PVvapour pressure of water. MI T2

Shaft horse power at ML2T3-PSt0 approach speed

PT .. port '

--

-static pressure at propeller M12 T3shaft

R propeller torque MI2T-22 -2

QAmachinery torque, astern MI T

Qt0machinery torque ahead at MI2T-2stop order

Q0 machinery torque, max 2T2continuous.continuous

radius of stopping track-

R ship hull resistance MLT2

R0R at approach speed MLT2-

r correlation factor

rs

SC

U-

V

V

W-

VRO

Symbol Definition Physical dimension

correlation factorfor eq (46)

correlation factorfor eq (47)

distance

head reach1:1

SLlateral reach

STtrack reach

Sy one standard error ofestimate

STc STcalculated

SB starboard

ST. steam turbine

propeller thrust

TApropeller thrust astern

astern thrust at ,dead inTAis the Water

Student's t

thrust deduction factor -

t time

tRtime.to reduce shaft, speed.or pitch to zero

Student's t for eq (46)

tastopping time

tt. StUdent'-s t for eq (47).

tscalculated_

MIT-2

MLT-2

MIT-2

-term of uncertainty MIT2

ship speed

approach speed LT-1

-1relative longitudinal wind LT

speed

absolute longitudinal wind 1JT-1

speedLT-1VR at stop order

wind resistance-2MLT

- 7 -

Symbol Definition Physical dimension

0

wake fraction -

x longitudinal distance betweencentres of origin of coordi-nate system and gravity

-

Yrforce-yaw velocity derivative MLT1

Y' Yr/0,5i3w171Tp2 d

Yvforce-sway velocity derivative MT-1

Y' Yv/0,5p,NVIppd

number of blades

a retardation factor

a mean value, m..v.amean

a due to turning,aC meanaa due to ship hullR "-mean .resistance,

a due to propeller thrust,mean McV.

a meana due to uncertain influences

U

mean a due to wind resistance, m.v.aW

.0 correction factor, eq (15)-2

A weight displacement MIT

V volume displacement3

1Sshafting efficiency

1R0hull efficiency at stop order

TIP()open water propellerefficiency at stop order

1R0relative rotative efficiencyat stop order

1tot toverall propulsive efficiency0 at stop order

de course stability numberduring crash stop,

V Pearson's coefficient _

of variation 1)

PAmass density of air MI-3

Pw mass density of water MI-3

W at stop order MLT-2

8

Symbol Definition Physical dimension

a'cavitation number at dead -pull astern, 2)

Y turning angle, yaw angle deg

statistical m.v. of T deg

Y angular Velocity, m.v. T-1

1) one standard deviation 100/arithmetic mean value

2)(P - Pv)/ 0,5 pw /12.11 D2

INTRODUCTION

The volume of transport by ship is expected to more than

double during the next ten years. There will be an increase

in the number of ships. HAINES 1971 estimates that 300-400

ships per day pass through the English Channel. For the

Uraga Channel, the entrance to Tokyo Bay, the number of

passages daily was 700 in 1968 and according to an EDITORIAL

1970 in Zosen, 50 ferryboat crossings of the main sea lane

should be added to this figure. HAINES 1971 reports 1300

vessels per day for the Inland Sea. Among other regions

with high density traffic are the straits of Malacca and

Singapore, the Persian Gulf, the Danish Straits, the St

Lawrence Seaway and the Panama Canal entrances.

The continuously existing requirement to improve the economy

of transport has resulted in an increase in the size of

ships and an increase in the speed for some types of ship.

The number of tankers above 200.000 dwt and high speed

container ships increases rapidly. As a result of the

building up of complete container handling systems at

large cost, the existing severe requirements to maintain

time schedules will be further accentuated. This can force

the master to maintain a higher speed in congested waters

than safe manoeuvring permits. According to HAINES 197111

every year approximately 7 % of all vessels above 500 grt

9

are involved in collisions. To this figure must be added

an allowance to cover groundings due to faulty manoeuvring

when the vessels are in the collision risk zone. It can

be estimated that 10 % of all vessels collide or run

aground. Naturally the majority of these casualties occur

in sea lanes with high density traffic, KOSTILAINEN 1971

and PORRICELLI et al 1971.

Recently there has been discussions in many places and

papers have been published on the question of the

manoeuvrability of large tankers especially having regard

for the risk of oil pollution. A case can be made for

giving special attention to large tankers but this way of

dealing with the problem of ship manoeuvrability is

insufficient. HAINES 1971 makes a survey of 40 collisions

in the Dover Strait. Of the 80 vessels involved 65 were

not tankers. The result of a collision between a tanker

and a general cargo liner or a ferry can be of the same

degree of severity as between two tankers. The risk of

loss of life is of most importance, pollution is a

secondary risk. Further, a ship which has run aground or

is party to a collision is usually involved in cost

consuming assistance and repairs. To this should be added

the cost of freight loss and increased insurance. Con-

sequently, it is of tha greatest importance to be able

to estimate the manoeuvring characteristics of different

ship types under varying conditions.

PAFFETT 1971 (I) gives a valuable general survey on

various aspects of safe manoeuvring, in which both the

human and technological influences are discussed. It is

necessary to improve the manoeuvring qualities of ships

and also to supply the master with reliable and simple

means to increase the knowledge of manoeuvring in

critical situations.

The decision to steer away from a risk zone or to carry

- 10 -

out a crash stop can only be made by the master dependent

on the circumstances prevailing at the moment when action

is required to be taken. To some extent TANI and FUJI

1970 deal with these questions. KENAN 1972 in an interest

ing approach investigates the manoeuvring qualities when

combining rudder and propeller for two Mariner, ships.

Thereby graphs of critical ranges for a number of vessel

manoeuvres are developed. However simplified assumptions

have been introduced such as for instance that the

propeller and rudder do not interact in a manner more

complicated than they do at 15 knots even at a full

astern manoeuvre. It is thus assumed that the rudder is

effective also when the propeller is reversed which is

not-supported in practice.

The stopping characteristics of ships will be considered

in this study.

MEANS AND METHODS OF IMPROVING THE STOPPING QUALITIES OF.. SHIPS

It is expected that in the future especially single screw

ships to an increasing degree will be designed to carry

equipment for improving their stopping qualities. Stopping

-methods by the use of rudder and propeller have been

investigated by among others JOURDAIN 1965, GROSSMANN 1971

and CLARKE et al 1972.. Multiple screw ships usually have

a better stopping performance than single screw ships.

Only the latter type will be discussed here.

To stop the ship more effectively, so called braking

flaps in the bow have been suggested by JAEGER 1963 and

JAEGER and JOURDAIN 1968. CLARKE and WELLMAN 1971 descripe

brake flaps placed in the afterbody of the ship. MITSUBISHI

1970 discuss the use of underwater parachutes to stop

large vessels. Mitsubishi have made tests in model, half

and full scale. DUPORT 1968 and ENGLISH 1968 have applied

for patents on Y-shaped canals in the bow. Duport suggeSts

an active duct i.e. including a pump, while English's

solution is a passive duct. The latter system is also

described in ENGLISH 1971. The above mentioned types of

manoeuvring devices are restricted insomuch as being

primarily efficient when proceeding at high speed. Further

they are relatively complicated to build into a vessel

or to operate. Although a relatively large number of

tests as well as theoretical studies have been made, the

basis for a practical solution is still considered to

be lacking. DICKSON 1971 is of the opinion that it is

not necessary to use this type of equipment. At NFL such

systems have been studied. On the basis of these findings

PAFFETT 1971 (II) states that: "The verdict in all cases

was that the devices undoubtedly worked, but the improve-

ments achievable would be relatively modest".

In order to reduce the time for shaft reversal for super-

'charged fourstroke diesel engines, extensive studies-

have been. .carried out. As a result a method for shortening

tt.e time for shaft reversal has been developed based on

the decompression method Klaunig's discussion of ILLIES

1969; °NISH' et al 1970.

CLARKE et al 1972 report from rudder cycling tests with

a 193.000 dwt turbine tanker. The rudder cycling programme

which BSRA conducted from an approach speed of 12,7 knots,

resulted in a track reach of 1.850 m compared to approxi-

mately 3.000 m at an ordinary crash stop manoeuvre with

the same approach speed. The method required five rudder

manoeuvres syncronized with five changes in shaft speed

and three alterations in the course of the vessel. Under

discussion with some shipowners, masters and shipyards

the method is considered to be complioated and as a

result there is a risk of making a faulty manoeuvre, see

also JOURDAIN 1965. The advantage of the method is that

the vessel is under steering control during the main

part of the retardation. pROSSMANN 1971 deals with rudder

cycling with an appreciably'smaller steam turbine vessel

and claims a reduction in stopping distance and time with

rudder cycling. However, the time for shaft reversal

during the ordinary crash stop tests was comparatively

- 12 -

long; 95 and 120 seconds. Until more tests are carried

out and have been evaluated, care must be taken when

comparing the method of rudder cycling with that of an

ordinary crash stop even assuming that the above mentioned

syncronization is managed successfully. A rudder cycling

test according to the BSRA method and an ordinary crash

stop test were carried out with a fully loaded motor

tanker of 150.000 dwt from an approach speed of 15 knots

The track reach with rudder cycling was 3.700 m and with

crash stop 2.600 m.

It is now generally accepted that a ship With controllable

pitch propeller (app) can have significantly better

stopping qualities than an equivalent ship with fixed

pitch propeller (FPP). Among others this has been stated

by GETZ and REFSNES 1958, HOOFT and VAN MANEN 1968,

ILLIES 1968, VAN GUNSTEREN 1970, NORRBY 1970, RITTERHOFF'

1970, INUBUSHI and SAKAI 1971, MASUYAMA 1970 and 1971, 1

OKAMOTO et al 1971 and AUCHER 1972. The reason for the

better stopping ability with the CPP vessel is that a

higher astern power can be used and utilized rapidly

after the stop order is given.

AVAILABLE DATA ON BOARD FOR ESTIMATING SHIPS' STOPPINGQUALITIES

Stopping tests are usually carried out in connection withthe trial trip. At some shipyards an appreciable time

for the stopping tests is reserved. Then crash stop from

full, half and slow speed ahead and a coasting test can

be included.

At the trial trip the stopping test is usually carried

out in non-restricted waters. Those engaged in the test

are well prepared. Consequently, the stopping order is

executed without loss of time and with due consideration

for the characteristics and condition of the propulsion'

machinery. The results thus added to the performance data

for the vessel are usually more favoUrable than what is

- 13 -

achieved in practice. Then the ship must be stopped in

the shortest possible time and distance in a critical

situation for which sometimes no preparations have been

made. This explains a statement made by CHURCH 1957:

"Even so, ships usually stop as well as can be expected,

except on occasions when collisions occur".

As a rule the ship is equipped with a plotted crash stop

track including a graph of ship speed and shaft speed

over time. For, a ship equipped with CPP the propeller

pitch is also included in the graph. Usually for a

tanker these results are for full load condition and for

a bUlk carrier or general cargo liner they are for some

specified ballast condition. This type of information

is difficult to interpret and use when differing condi-

tions of displacement, trim and approach speed appertain.

A more practical graph has been suggested by Oosterveld

in the discussion of TANI 1968 and is also shown in

SHELL 1968. TANI 1969 also uses 'this type of graph. IMCO

have recommended that such a diagram, shown in figure 1,

be added to the manoeuvring data for "large ve6sels and

those carrying bulk chemicals", PRICE 1970. The majority

of crash stop tests have been carried out at full

astern shaft speed. Consequently, the diagram, based

mainly on crash stopping tests, is not so accurate at

lower shaft speeds and care must be taken when estimating

the performance under those conditions. Further the

diagrams which have been presented only deal with full

draught and these should be complemented by diagrams

under part load conditions. In SHELL 1968 the graphs

are based upon the mean stopping values for each class

of tankers. From the safety point of view it would be

better to supply the ship with a diagram showing a so

called risk distance and risk time at a crash stop.

It is also valuable for the master to be able to estimate

the probability of the vessel turning to starboard (S13)

or to port (PT) during a stopping manoeuvre. The general

14 -7

opinion is that the vessel turns to SB but there is no

stated degree of probability for turning SB versus PT.

There are different ideas of how a ship with CPP reacts

during a stopping manoeuvre as regards turning.

METHODS Fait. CALGULATING STOPPING PERFORMANCE

It is necessary to clarify how the results of a stopping

test should be appraised. The scatter in performance is

large and sometimes difficult to explain. The appraisal

is comPlicated by the influence arising from turning

forces, suction effects, wake variation, cavitation,

air drawing, oblique flow to the propeller and hull

resistance during retardation, DAILY and HANKEY 1953,BINDEL and GARGUET 1962, JAEGER and JOURDAIN 1962,SAUNDERS 1965, HARVALD 1967 and LOVER 1969. The picture

is rendered still more complex due to the transient

nature of the stopping manoeuvre.

A large number of stationary and quasistationary methods'

for calculating track reach and stopping time have been

developed for ships with FPP based on relatively simple

assumptions, DRESSLER 1912, ROBINSON 1916, NORDSTROM

1931, ROBINSON 1938, HORNE 1945, CHASE and RUIZ 1951,HEWINS and RUIZ 1954, CHASE et al 1957, MINIOVICH 1960,LINDGREN and NORRBIN 1962, TREFETHEN 1962, SAINSBURY

1963, BOATWRIGHT and TURNER 1965, CRANE 1966, HARA et al

1966, TANI 1966, CRANE 1967, HARVALD 1967, TANI and

ISHIKURA 1967, GOODWIN et al 1968, GROBE 1968, ILLIES

et al 1970, RUMS 1970, TANI and ENOKIDA 1970,

CLLRandWELLMAN 1971 and EHRICHE and GROSSMANN 1971. In all

these methods it is assumed that the course of the vessel

is constant during the stopping manoeuvre, the propeller

characteristics from tests at atmospheric pressure are

assumed valid with no allowance for the influence from

cavitation or air drawing and with the exception for

HARVALD 1967 it is assumed that the suction and wake

factors are constant. In spite of these simplifications,

-215-

and of the fact that the methods differ somewhat one from

another, several authors claim good conformity with results

from actual crash stop manoeuvres.

For ships with CPP also methods have been suggested to

calculate track reach and stopping time, CHASE et al 1957,

KITO 1959, MARTIROSOV 1962, GOODWIN et al 1968 and OKAMOTO

et al 1971. The method used for calculation is the same

as that for ships with FPP with the exception that at

stalled conditions a CPP has other characteristics than

an FPP.

COLLECTION AND PROCESSING OF CRASH STOP TEST RESULTS

In order to shed more light over the stopping charac-

teristics of ships, a collection of crash stop results

and ship data has been made with the aid of shipyards,

shipowners, universities, private persons and literature.

Test results from the following references in the

literature have been used: HEWINS et al 1957, GETZ and

REFSNES 1958, HEBECKER 1961, WILSE 1962, MOCKEL and

HATTENDORFF 1966, PARKER et al 1966, HARA et al 1966,

CHIHAYA 1967, HOOFT and VAN MANEN 1968, JAEGER and

JOURDAIN 1968, JOURDAIN 1969, JOURDAIN and PAGE 1969,

SCHIELE et al 1969 and .editorials from HANSA 1962,

SHIPPING WORLD AND SHIPBUILDER 1967, SHIPPING WORLD

AND SHIPBUILDER 1969, SWEDISH SHIPPING GAZETTE 1969 and

ZOSEN 1969. For every test a table has been filled in.

An example is shown in table 1. The test results are

from 25 shipyards in Denmark, England, France, Germany,

Japan, Norway, Sweden and USA. In total 330 crash stop

tests in ballast and full load have been collected,

of which 255 are for ships with FPP and 75 for ships

with CPP of KaMeWa design. The distribution of the

tests is shown in figure 2 on the basis of the deadweight

tonnage. The material has been collected without reference

to any special size or' type of ship. The purpose has been

- 16

to cover the whole range of single screw ships. As the

tests originate from a large number of shipyards in

various countries, it can be assumed that they constitute

a representative average and therefore the material

should not be influenced to any great extent by the

special practices of any single shipyard. 60 % of the

tests have been carried out with vessels with a dead-

weight above 20.000 tons.

Some Of. the tests are incomplete. In these, instances

only part results have been used in a wider context.

FUrther it has been possible to exclude some obvious

faults made during Certain particular tests When this

data has been plotted.together With the rest of'the

material. The number of tests included in the various

parts of the study is .stated in each case

Track reach ST is the distance. along the stopping track.

It has been calculated as the area below the speed-time

curve. Thus ST is based on speed log recordings and not

on fixed plotting from for instance, Decca. In this

instance the former method is to be preferred as ST

is then not influenced by current. The ST value is the

distance travelled through the water. The accuracy of

the log method is + 0,5 - 1 % down to 4 - 5 knots. At

lower speeds the accuracy is less; + 1 - 4 % due to

disturbances from propeller wash and oblique flow in

the last period of the manoeuvre. However, the recording

of the track reach is only influenced by this to a

small degree. The stopping time has been taken from

fixed plotting which is more accurate than speed log

recording at low speeds.

Head reach, SH, is the distance measured along the

ship's initial course from the bow to the point in the

stopping track which is most distant in this direction.

Lateral reach, SL, is the distance from the ship's centre

line measured perpendicular to the initial course to the

With very few

in open water

by the effect

the few tests

- 17 -

most distant point on the Ship during the manoeuvre,

see figure .3 for definitions.

Practically in all tests the rudder has been kept amid-

ships in accordance with the usual practice. In a few

tests the rudder has been used but this has not influenced

the result. A general observation is that during an

ordinary crash stop the ship cannot be steered due to the

very low or no rudder effectiveness, AUCHER 1972.

A crash stop manoeuvre is characterized in that the

reversing of the propeller shaft or with CPP, reversing

of the propeller pitch, is carried out in the shortest

possible time. This usually means within 15 - 45 seconds

for FPP, diesel engine (DE) and 30 - 75 seconds for FPP,

steam turbine (ST). The corresponding figure for CPP is

approximately 15 seconds. After the reversal the highest

possible astern shaft speed is used. For the collected

material on FPP DE this means an astern shaft speed of

approximately 65 % of the full ahead value and for

FPP ST 50 - 55 %. With CPP between 90 and 100 % is

utilized when stopping.

exceptions the tests have been carried out

whePe,h/d>4. Thus they are not influenced

of shallow water, FUJINO 1969. When plotting

where h/d7. 2 these fall well within the

mean values of the remainder of the observations and

have consequently been included in the study.

The design power used is the maximum continuous rating

of the machinery. If the CPP ship is equipped with

auxiliary shaft driven devices the stopping test has

been carried out with these working. This means that

the auxiliary power is deducted from the power of the

main machinery. It has been estimated that the auxiliary

power used during the test is 70 % of the design power

of the auxiliary machinery.

18

All FPP are righthand rotating while for CPP both right

and lefthand rotation exists. The reason why a CPP is

frequently supplied lefthanded is that thereby it will

act on the vessel with the same force direction during

a stopping manoeuvre as an FPP. The reason why lefthand

rotation is not always specified is that the majority

of main propulsion machineries are built for righthand

rotation.

In the purpose of the study of the various combinations,

it has been found convenient to divide the statistical

material from the stopping tests into three.groUpa:

Fixed pitch propeller, diesel engine (FPP DE)

Fixed pitch propeller, steam turbine (FPP ST)

Controllable pitch propeller, (CPP DE ST)diesel engine and steam turbine

Each group consists of tests with ballasted as well

as fully loaded ships and also with approach speeds

of design value and down to 3 knots.

So far as is possible the material has been converted

into nondimensional form in order to eliminate the

influence of size to some extent. Finally the material

has been dealt with and appraised by simple statistical

methods described in MORONEY 1962 and EZEKIEL and FOX

1967.

COURSE STABILITY DURING_THE STOPPING MANOEUVRE

Before considering the stopping performance of ships,

the shape of the stopping track and its depehdence

on the ship's course stability at a stopping manoeuvre

will be studied.

It is a matter of general knowledge that a fully.

loaded vessel sheers more during a stopping manoeuvre

than a ballasted vessel. During such a manoeuvre when

- 19 -

the propeller thrust is reversed the effect of the

rudder is eliminated. This occurs in the early stages

of the manoeuvre, i.e. before the reversing of the

shaft or the pitch. HOOFT 1970 indicates that the rudder

effect is decreased when the speed of the propeller jet

is reduced and is zero after the reversal. Based on PMM

tests with an 8 m model of a 190.000 dwt tanker in full

load it is confirmed by WAGNER SMITT and CHISLETT 1972

that the rudder efficiency is negligible during a

stopping manoeuvre.

When the propeller drives the ship ahead a force in the

stern of the ship perpendicular to its longitudinal

axis is developed due to the local loading of the

propeller. This is here called the propeller rotation

effect and is described in detail by SAUNDERS 1957.

For a righthanded propeller this results in the side

force being directed towards SB. According to Norrbin's

discussion of CLARKE and WELLMAN 1971 it is some

3 - 6 % of the ahead propeller thrust. The side force

directed to SB gives the ship a tendency to sheer to

PT. When steering on a straight course the side force

is compensated for by some degrees of SB rudder angle.

For the sake of simplicity it has been assumed that

the influence of wind, waves and current is negligible

here.

During ,a stopping manoeuvre with an FPP the side force

on the stern will change direction due to the reversal

of the propeller thrust. The stern is then forced to

PT and as the rudder has no effect in this condition

the vessel usually sheers to SB. On the other hand

for a vessel with OPP no side force reversal takes

place when the propeller thrust is reversed. From the

model tests by WAGNER SMITT and CHISIETT 1972 it can

be estimated that the side force from the propeller

on the hull att full astern shaft speed is in the order

of 10 - 20 % of the ahead design speed propeller thrust

for this vessel.

- 20-

With a lefthanded CPP when the pitch is reversed, the

direction of the side force is the same as for a right-

handed FPP. For the remainder of this study a ship with

a lefthanded CPP which sheers to SB will be designated

a SB sheer. The same is valid for a ship with a right-

handed CPP sheering to PT. If these ships sheer to the

opposite side they will be designated PT sheering.

Usually with a CPP the thrust is reversed in a consider-

ably shorter time than with an FPP and consequently the

influence on the sheering of the vessel due to the

propeller rotation effect will appear at an earlier stage.

The sheering is not always directed to SB. The influence

of outer forces on the ship results in some instances

in sheers to PT. This is the case especially at, or

close to, the full load condition when the ship usually

is directionally unstable during a crash stop. It is then

more sensitive to disturbances from wind and current.

Also extensive air drawing and production of eddies at

the manoeuvre can influence the direction and amount

of sheer, BINDEL and GARGUET 1962. Full scale observa-

tions have shown that air drawing exists whether the

propeller is partly or completely submerged.

However, air drawing has only secondary influence on

the direction of the sheer. The essential is the

influence of external forces mainly due to wind and

eddy formation. and internal forces on the vessel during

the first phase of the stopping manoeuvre when the

rudder effect is lost. These forces directly in4luence

the direction of the sheer. In the second phase

propeller rotation effect dominates so that an alrqady

initiated SB sheer is increased while the PT sheer is

decreased.

In ballast., when the vessel as a rule i8 stable on

course, it is less sensitive to outer disturbances.

In the majority of cases the vessel then sheers SB,

21

i.e. in the second phase the propeller rotation effect

again dominates.

A"COURSE STABILITY CRITERION FOR STOPPING'MANOEUVRES'

The difference in the sheering tendency between a fully

loaded and a ballasted vessel whilst stopping is to be

further investigated having regard for course stability.

The criterion for dynamic stability in straight line

motion for a vessel with its rudder in the neutral

position can be written

Y (N CB BMd2

xGV) - Nvr - pw CB Bm d2 V)> 0 (1)

The deduction is shown by a.o. ABKOWITZ 1964 and further

work to obtain a practical criterion has been carried out,

WAGNER SMITT 1970 and 1971. Based on 55 PMM tests with

35 different models varying from trawlers to large tankers

force and moment coefficients Y;., Y,,N; and N. have been

evolved as a function of d/Iipp. Using these values in

the above.expression (1) Wagner Smitt writes the criterion

as

41 3,88C Bu L /( L ) 1.5'23

+ 0,0050 ( ) >0pp

Wagner Smitt states that the criterion is not valid

for trimmed vessels and shows that approximately there

is a linear relation between the force and moment

coefficients and the square of the aspect ratio i.e.

(d/LF2)2. When test results with trimmed models are

included the scatter is relatively large.

A more correct expression for aspect ratio is 2d2max/

lateral area. Assuming that d =dmax A

the expression

for aspect ratio can approximately be written

44A/[Iippi(dF/dA) + . To adapt the criterion to

(2)

br....,-"....-4Vreammrom.,.."....

- 22-

trimmed vessels the force and moment coefficients are

plotted as a function of the latter expression for aspect

ratio. Then the expression's for force and' moment

coefficients will be of the type

[v - YvL +

dF

pp dA

dA 2

WhereC is constant. -Thus the inequality (2) can be

written v

LV PPL72.2)3

< Ci d 2A. Ad + 1)

A

Lacking coefficients for trimmed vessels Wagner Smitt's

values for vessels on even keel have been used here.

This gives

5,4L L 3

<, V + 0,0013 (U)dF 2 dA

dA3.

(cr + 1)A

The conditions for dynamic stability during a stopping

manoeuvre &re considered to be more complicated than for

steady propulsion ahead. To some extent this question

has been dealt with by NORRBIN 1964 and HOOFT 1970.

WAGNER SMITT and CHISLETT 1972 show that from the above

mentioned PMM tests the dynamic stability during a crash

stop Is closely the same as for steady ahead propulsion

and uninfluenced by the speed of propeller or ship.

Howev-er, the question requires further scrutiny because

with the existing material it is difficult to make

general forecasts regarding sheering tendencies during

stopping manoeuvres.

The inequality (5) has been used to further' studythe

tendency to sheer when stopping. If Wd3A is' less than

the value' in (5) thenthe vessel is stable as usually

defined for steady propulsion ahead. From 'the. stopping

(3)

(5)V

"fir

23

, 1/3tests SL/ V and Y for sheer SB and PT have been

plotted as a function of (V/d3A)/(v/d3A)c where index

C indicates a critical value based on inequality (5).

This has also been done for sheer against and with the

wind, see figures 4, 5, 6 and 7. It is evident that a

definite boundary value exists for (V/d3A)/(V/d3A)c.

For (N7/d3A)/(V/d3A)c < 1,1 the vessels sheer moderate-

ly while for values > 1,1 the sheer and its scatter is

appreciably greater. Obviously the dynamic stability

during a crash stop has a critical value around 1,1.

A course stability number is now introduced. The

critical stability value during a crash stop is for

simplicity put at 6f = 1, i.e. where (V/d3A)/(V/d3A)c

= 1,1. Based on expression (5) and WAGNER SMITT and

CHISLETT 1972 a criterion for dynamic stability during

a stopping manoeuvre can be written thus

5 9 L L 3

(34)dF

PP + 0,0014 (--P-2)2 dA

(6)d' CA

A d, (7 1),A

This is shown in figure 8. In order that the vessel shall

be dynamically stable during a crash stop its V/d3A

must be less than the value in equation (6). If for

example a ship has dp/dA = 0,95 and Lpp/dA = 18 then

(V/d3.)= 36. The actual value for V/d3A is 39A CA

which means that the ship must be considered to be

dynamically unstable during a crash stop; the value

is > 1. By dividing the material from the crash stop

tests into two parts, ee>1 and dlo<1 a Clearer picture

of the sheering tendencies is obtained.

SHEERING TENDENCIES

In order to study with what probability a vessel sheers

to SB or PT, with or without the influence of wind,

figures 4, 5, 6 and 7 have been investigated. To this

material has been added results from stopping manoeuvres

24

where the direction of sheer and relative wind are known

but not SL and LP . The result is divided into the two

groups 4f > 1 and g < 1 and is shown in figures 9, 10

and 11. To simplify the comparison in figures 9 to 11 it

has been assumed that for ahead propulsion all FPP are

right hand rotating and all CPP are left hand rotating.

As mentioned above this implies that the side thrust

on the vessel from the reversed CPP then has the sae

direction as from the reversed FPP. Thus if the CPP is

right hand rotating the numbers for SB and PT are to

change places in figures 9 to 11. The windward side

must also be-changed.

SHEER WIN, TYPE OF 'PROPELLER, MACHINERY AND PROPELLERROTATION EFFECT ARE TAKEN INTO CONSIDERATION BUT.EXCLUDING INFLUENCE: OF wIya

For df>1 figure 9 shows SB sheer dominant with approxi-

mately 60 % for FPP DE and ST and with 68 % for OPP.

The tendency mentioned above for SB sheer has further

Increased for H < 1. There the SB sheer accounts for

70 % with PPP ST, 78 % with FPP DE and 82 % with CPP.

The explanation is that the directionally stable ship

maintains roughly the initial course during the first

phase of the stopping manoeuvre and in the second phase

when the speed is low is influenced principally by

the propeller rotation effect which tends to turn the

ship to SB. Further the influence from propeller and

machinery is larger when the ship is in ballast than

in full load.

SHEER WHEN TYPE OF PROPELLER, MACHINERY, PROPELLERROTATION EFFECT AND WIND DIRECTION ARE TAKEN INTO

CONSIDERATION

With of>1 and the wind on the SB side approximately

75 % of the ships with FPP turn SB. The corresponding

value for CPP ships is 80 %, however care must be

- 25 -

taken with this value because the number of observations,

is only 5. With the wind from PT, ships with FPP ST..and

DIE turn SB, 40 % and 50 %respectively. The CPP ship falls

Off from the wind and sheers SB in 65 % of all cases.

Evidently the sheer of the CPP ship is mainly influenced

by the direction of the rotation of the propeller and

not by the direction of the wind wheh it is on the PT

side. For FPP ST and Eg the influence of the PT wind is

strongest. With the wind on the SB side the effect of

wind and propeller rotation add up resulting in an

increase in the sheerability to SB.

With al?<1 and SB wind, 75 % and 86 % turn SB for FPP ST

and DE respectively. For CPP the value was expected to

be between 80 % and 90 % but in this study is not more

than 57 %, probably due to the study being based on 7

observations only and in consequence not being reliable.

The tendency to luff into or fall off from the wind can

be noted in, for instance, SHEARER and LYNN 1960 or

GOULD 1967. There it is shown that vessels on even keel

with the superstructure aft have a positive wind moment,

i.e. they have a tendency to luff for a total sector

of a relative wind direction of 220° - 280° from astern.

For ballasted vessels trimmed on the stern and for vessels

with the superstructure midships or forward the "luffing"

sector is smaller, approximately 180°.

This survey of the sheering tendencies can be used for

estimating approximately the probability that a vessel

will sheer in a certain direction when making a crash

stop manoeuvre. The values for sheer given in figures

9 to 11 indicate the sheering probability. By use of

the multiplication law the probability may be estimated

as to what extent two ships meeting on counter course

will sheer against each other for instance.

- 26-

Assume that two ships on counter course meet in a

passage and keep on the SB side. One of the ships is

a fully loaded tanker (0 >1) with FPP ST and the other

is a general cargo liner in ballast (0 < 1) with FPP DE.

The tanker has the wind from PT and the liner from SB.

The probability that both ships sheer PT in a stopping

manoeuvre is 0,60. 0,14 = 0,08. This means that in 8

cases out of 100 the ships can sheer into the centre of

the passage thereby increasing the risk of collision.

If the two vessels steam the passage on the wrong side

(PT) which can occur, the risk that both. sheer SB

during a stopping manoeuvre is Markedly increased;

0,40 0,86 = 0,34. In this particular case the risk that

the vessel will Sheer against each other is approximately

four times as great.

Generally it Can be concluded for the. material that

if two vessels on counter, course steam a passage

keeping'SB the probability that both will sheer. PT

does not exceed 0,20. If the vessels steam the PT side

of the passage when making a.crash stop the corresponding

value for both vessels sheering SB is up to 0,67.

THE TURNING ANGLE

From the view point of handling the ship it is of

interest to know the final turning angle y the vessel

will reach during a stopping manoeuvre. In figure 12

the mean turning angle for FPP and CPP at 0> and < 1

is shown with approach speeds above 13 knots. Sectors

for + one standard deviation have been drawn approxi-

mately. Above 13 knots y is not to any measurable

degree dependent on speed or ship size while for lower

approach speeds V decreases markedly. As there is no

significant difference between the turning angle for

FPP DE and ST these results are presented together in

the figure.

- 27 -

At FPP >1 the difference in LP-SB and PT is not

significant. On the other hand the difference in

turning angle SB and PT is highly significant for

FPP,dT < 1. From the plotting of the course during

the stopping manoeuvre it is seen that in the second

phase the vessel turns SB due to the propeller rotation

effect whether its final sheer is SB or PT. This

explains the difference SB and PT in the mean turning

angle i-7 which is 13° for df >1 and 29° for g < 1.

For CPP, d-p >1, SB sheer Ti) is 133° and its standard

deviation is 24°. The .corresponding values for FPP are

125° and 49°. The smaller scatter for the vessel with

CPP is explained by the fact that the crash stop as a

rule is performed more uniformly than for a vessel

with FPP. The difference 8° in mean turning angle is

not significant. At oe< 1, SB sheer the turning angle

and scatter are approximately equal for FPP and CPP.

For FPP as well as for CPP there is a highly significant

difference for 4) between d-P > and < 1.

DETERMINING TRACK REACH AND -STOPPING TIME

During a crash stop manoeuvre the vessel along its

longitudinal axis is mainly retarded by the ordinary

hull resistance, turning resistance, wind resistance

and the braking force of the propeller. These forces can

be estimated approximately as a function of time. However,

as mentioned above the quantitative knowledge of the

physical basis is insufficient. As a consequence the

calculated results sometimes can differ appreciably from

the real crash stop results. This is so even when using

a complete system of equations to solve the crash stop

distance and time.

Ah approximate solution to the complete system of stopping

equations can be made by assuming that the track reach

-

and time can be derived from

A dV- + kx = -F

F. can be expressed as

F .1/A-C+W+ TA-(1- 7 t).. (8)

For ordinary merchant ships this is considered

acceptable.

The qualitatively and quantitatively unknown influences

during the manoeuvre are here represented by a term

of uncertainty, the force U which should be added to

the right side of equation .(8).

The purpose here is to arrive at simple relations

between the total braking force and the stopping

distance as well as the stopping time. It is therefore

considered sufficient to investigate the mean value

of the braking force Fmeanwhich can be expressed as

-

F + C + W + TA (1 - t) + UImean-mean

THE RETARDATION FACTOR

The retardation factor a is here defined by

1 dVa =g at

. or

- z1(1

F.+ kx)

From figure 13 and equation (10)

( V dVV dt g a

(7)

0

(12)

- 29-

a is an unknown function and this complicates the

solution of the integral. Study of the V-t curves

in the stopping material suggests that the

retardation is roughly constant during the manoeuvre.

This simplification is introduced, thus

cc = amean

Then from equations (12) and (13) the uncorrected

track reach is obtained2

VOS (uncorrected) = -T 2 g amean

By introducing the factor p in equation (14) an

approximate relationship is obtained. p corrects for

the arbitrary shape of the V-t curve, see figure 13.

This implies that p expresses the area ratio

ST2

t Vs 0

consequently with equations (14) and (15) the non-

dimensional expression for track reach is

g ST

Vo2 atean

In the collected crash stop results the meah-ValUe.

of p is 1,09 fOr EPP 3, 1,04 for FPP ST and

1,02 for CPP. The .standard deviation is approximately

0,1. A simple but rough estimate of the stopping track

can thus be made when approach speed and stopping time

are known. A Corresponding 0 value from HARA et al -

1966 is 1,04 for 106 tests of which 75 are with FPP $T

and the rest with FPP DE.

For the stopping time a non-dimensional expression is

obtained from equations (10) and (13)

g ts1 (17)Va

0 mean.

The mean value of the retardation factor a canmean

(i3)

(14)

By assuming that the deceleration is linear with time

and that the hull resistance can be expressed as

RVn (20)

where n is constant through the speed range the following

relation is valid

R0Rmean n + 1

is the hull resistance at approach speed.

_

PTit t t - 'PO T/R0 T1E10 TISO

St0 -

Thus

aR mean-

2 n '/tot ton + 1

- 30--

nt0 0

A(1 + kx)(23)

It .should be noted that n is an approximate mean value

for the whole resistance curve. Further the estimation

9f TItot t can be rather rough especially at off design0

conditions.

(18)

be written with equation (9) as

amean= - mean - mean - mean [TA(1-tAmean UmeanA(1+k ) A(l+k ) (1+k) (1+k) A(1+k

or

amean R mean+

aC mean W mean+a

T mean+ a= a + a

II mean(19)

A review of the various retardation factors is made

below. It will also be shown that for practical

estimation of crash stop performance only factors in

have to be used.aT mean

:RETARDATION DUE TO HULL RESISTANCE.

mean

. where 'P meanis the mean value of the angular velocity.

Finally093 L + k) Lj

meang (1 +.kx)es,y

12mean (26)

The calculation of inertia or "added mass" coefficients

for a prolate ellipsoid moving along its major and minor

axis in potential flow has been made, LAMB 1945. This

gives kx and k values of 3 to 4 % and 92 to 94 %

respectively on the assumption that the ratio of the

major to the minor axis is in the same region as

Lpp/Bm for merchant ships i.e. between 6 and 8.

g R

R is the variable radius of the stopping track. The mean

value of C can be written

0,3LAP

A(1 + ky)`-.P2mean

2

-31 -

RETARDATION DUE TO TURNING RESISTANCE

During the stopping manoeuvre the vessel sheers and a

centrifugal force develops. Its longitudinal component

C contributes to the retardation. The derivation of C

can be made from figure 14. It has been assumed that

the mean value of the distance between the vessel's

centre of gravity CG and pivoting point CP is 0,3 Lpp

and also that the drift angle is small, HARA et al

1966; The position of the pivoting point for steady

turning at ahead propulsion is dealt with by NORRBIN

1971. There it is shown that the steady state pivoting

point is further ahead for a full than for a slender

ship and varies from 0,5 I of the centre of

gravity for a full tanker to 0,3 Lpp for a destroyer.

However, for the unsteady state the value of 0,3 Lpp

is chosen because of the accentuated turning towards

the end of the stopping manoeuvre. Thus0,3 L A(1 + k) V

CPP

- 32 -

For ordinary ships forms kx and ky can for instance _be

obtained from the results of extensive model tests and

studies by MOTORA 1960. The graphs are also shown in HARA

et al 1966, when Ii./BM, d'/BM and CB are known. Motora

mentions in a letter 1970 that the accuracy when estimating

kx and k based on model tests is approximately + 10 %.

This is estimated from collision model tests, MOTORA et

al1969.Thecalculationofkxand k with Motora's

method for the material investigated has resulted in

values ranging between 2,5 and 7,5 % and 60 and 80 %.

The values depend on the type of ship and whether it is

in ballast or full load.

Although the value of kx is not so important for the

solution of the stopping manoeuvre a number of values

are used by different authors. FROUDE 1874 was the first

to treat the subject. Then, among others, KEMPF 1928,

LEWIS 1929, VON DEN STEINEN 1933 and SMITH 1955 have made

contributions. The kx value for these investigations when

considering ordinary merchant ship forms is in the region

of 5 to 10 %. HARVAID 1967 suggests 5 %. Several authors

recommend 8 % "added mass" when calculating the stopping

results based on the tests by VON DEN STEINEN 1933.

HOOFT and VAN MANEN 1968 and OKAMOTO et al 1971 use a

figure of 10 %. Model tests at NSMB indicate that kx is

very markedly dependent on ship speed as well as

propeller shaft speed, HOOFT and VAN MANEN 1968 and

JAEGER and JOURDAIN 1968. The relationships are complicated

and further work is required before the results can be

used more generally. These findings are contradicted by

results from recent tests with models of large tankers

at Bassin d'Essais des Carnes de Paris and HyA,

Copenhagen. The first tests indicate a kx value between

5 % and 10 % whatever the ship speed or propeller

revolutions may be, AUCHER 1972. The latter tests with

a model of a 190.000 dwt tanker reported by WAGNER SMITT

and LANDSBURG 1972 show only a slight increase in kx

with ship speed. From these tests it can be estimated

The component of the relative wind alongships varies

during the stopping manoeuvre. This is due to the fact

that the speed of the ship decreases and also that the

ship sheers. The wind component W can be expressed as

2W = 0,5 cx pA VR A

Further

ts0,5 cx PA A rwmean t Ji VR2 dt

S 0

Whether Wmeanhas positive or negative sign, i.e.

contributes to shorten or lengthen the stopping manoeuvre

depends on the absolute value of the wind velocity and

direction, the vessel's approach speed and the shape of

the stopping track. It has been assumed that the direction

of the absolute wind is on the bow, that the vessel does

not sheer during the stopping manoeuvre and that the

deceleration is linear with time. Then the following

relation is valid

- 33 -

that the mean value of kx during a crash stop manoeuvre

is approximately 9 - 10 % for a tanker of 190.000 dwt.

Calculation of aC mean/ ameanfor the statistical

material shows, that the part of the total retardation

due to the turning can be as high as 30 to 40 % for

> 1, while the figure is lower for 4P < 1, not

exceeding 15 %. The influence of the vessel's turning

on the stopping result varies appreciably, especially

when e > 1. It is impossible to predict this influence

even approximately without knowing the shape of the

stopping track. On the other hand when analysing a

stopping manoeuvre where the track as"well as the course

have been plotted, the influence of the turning can be

estimated, see HARA et al 1966.

RETARDATION DUE TO WIND RESISTANCE

(27)

( 28 )

mean

9V- V

(W

2+

VI + 1)

VitoVR0

- 34 -

(29)

With equations (28) and (30) as well as data from

GOULD 1967 the retardation due to the wind resistance

has been estimated for a 10.000 dwt general cargo liner,

a 56.000 dwt ore-oil carrier and three tankers ranging

from 115.000 to 210.000 dwt in ballast as well as at

full load. This study shows that aW meanmean on a/ astraight course is essentially the same for the vessels

investigated.

In figure 15aW mean/ amean is presented as a function

of the absolute wind velocity at a ship speed of 16

knots for ballast and full load. Only in 5 % of all

cases aW mean mean/ a exceeds 0,09 and 0,15 in full

load and ballast respectively. This is based on wind

speed statistics from 179 stopping tests.

The values on influence alongships from wind on the bow

obtained from figure 15 are the highest which can develop

whilst the ship will sheer from its original course

during the manoeuvre. This decreases cx as well as VR.

Consequently in practice amean is only of minorWimportance for the total retardation during stopping

manoeuvres and the factor will therefore not be considered

further in this connection.

RETARDATION DUE TO THE STQUING ENERGY OF'THE 'PROPELLER

During the stopping manoeuvre the propeller thrust changes

sign in a relatively short time, before the shaft or pitch

reversal. The characteristic of a crash stop manoeuvre is

that the thrust vector rapidly changes direction and does

not decrease to any large extent in absolute value. The

thrust then usually stabilizes itself at a somewhat

reduced value, close to the dead in the water pull astern

- 35 -

value. The thrust revereal is carried out under non-

stationary conditions and the propeller blade sections

then work at extreme angles of attack and can thereby

cavitate heavily and also draw air. As a result the

rate of the reversal might have to be reduced to prevent

overspeeding or overloading the propulsion machinery.

With a CPP it is possible to keep a higher rate of

reversal provided that the propeller is equipped with an

automatic load control so that shaft speed as well as

machinery load can be kept within prescribed values. A

number of full scale observations of ships in full load

as well as ballast, indicate that the propellers (FPP

and UP) draw air especially during and just after the

reversal. Recordings of shaft speed and torque on ships

with FPP show large variations, HEWINS et al 1950,

HARA et al 1966.

Further, the thrust deduction factor varies as may be

expected, due to the fact that during the stopping

manoeuvre the propeller jet is stopped and then reversed

against the stern. The relations are complicated and have

been discussed by ODENBERG 1945 and HARVALD 1967 for

instance. In the latter report also the wake fluctuation

during the manoeuvre is dealt with. When calculating the

influence of the propeller on the stopping performance

it is common to simplify the problem by assuming that t

and w have the same values as for stationary conditions

ahead. TANI 1968 puts t = 0.

Several authors suggest that TA can be exchanged by TA

i.e. the bollard pullastern value ats

the

vessel has been brought to a,stop. NUterous.full scale

thrust measurements show that this is 4 reasonable

assumption.

From equations (18) and (19)[TA (1 -a= -T meanA (1 + icc) (30)

A

CTP 2 2 2pW nA

DPA

C p 2QA

3pw nA D PA

From this the wellknown relationship is obtained, CHASE

et al 1957

oPAA CTp--

) = CACQP A

-Assume that during the stopping manoeuvre the follOWimg

relation is approximately valid

LTA (1 - ci TAt (1 - t) (34)mean

c1 is a constant the mean value of Which according to

TANI 1968 is 0,925 for tankers. From (30), (33) and

(34) is thus obtained

c1t)

QA CAa -mean PA L(1 + kx)

RETARDATION DUE TO UNCERTAIN INFLUENCES

From the above it is obvious that conditions around the

propeller and stern are particularly complicated during

the stopping manoeuvre. To this should be added influences

from ship motions and retardation effects in the boundary

layer and in the ship waves. In the expression for aU mean

these influences should be taken into consideration.

The sign of aU meanis determined by if the mentioned

corrections are positive or negative. As knowledge in

this field is very restricted it is only possible here

to write

.meanaU mean -(i+ kx)

- 36-

For bollard pull astern the thrust and torque coeffiCients

.can be written, HEWIUS et al 1950 -

(31)

( 32 )

(33)

( 35 )

( 36 )

- 37 -

USE OF THE RETARDATION FACTOR

It is thus possible to state approximately the influence

of the various factors on the total retardation and

thereby estimate approximately the stopping performance.

Before equation (19) can be given a more satisfactory

solution the factors in au also have to be penetrated.1

Rough estimates however show that the dominating

influences in a stopping manoeuvre are obtained

froma T' R

a and ac From the above it is evident that

neither ac nor are possible to quantify with an

acceptable accuracy when predicting stopping performance.

For ordinary merchant vessels the quotient aT/ amean R mean

is roughly between 2 and 4 at a crash stop. Thus aT

is dominating.

In order to investigate the influence of different

_propulsion systems on the stopping performance this has

been studied as a function of factors in a T. Here

(1 t) and kx in equation (35) have been omitted.

in equation (16) is assumed to be constant within

each group FPI' DE, FPP ST and CPP DE ST and has

consequently been omitted. The following approximate

relations have resulted:

g S T PA= f -1

V2 QA CA0

g t PAP

Vf (2 "

0 CIA CA

The relationhilps (37) and (38) include the So Called

thrust constant CA from equation (33). The remaining

factors in (37) and (38) are relatively simple to

estimate.

It has been assumed that the torque can be approxi-

mately expressed by

nA

qA = % (1 - 0,4)0

38

(FPP DE and. CPP DE ST)

(FPP ST)

PROPELLER CHARACTERISTICS DURING, CRASH STOP

The CA value for FPP at bollard pull astern is set at

5,2 by HEWINS et al 1950 for the "Esso Suez". This

reference also states that the value conforms well

with the open water model test results in atmospheric

conditions reported by NORDSTROM 1948. HARVALD 1967

reports on tests carried out in atmospheric conditions

with two models of the Nordstrom propellers with

P/D = 0,6 and 1,2 for a bulk carrier and a trawler.

The bollard pull astern values in behind and open

condition are practically equal for the respective

propellers.

According to Perring's discussion of CONN 1932 and TANI

1966 there is a relation between CA and PA/D. A number

of open water model test results with merchant ship

types of FPP and CPP in atmospheric conditions have

been plotted in figure 16 based on material from CONN

1932, NORDSTROM 1945 and 1948, MORGAN 1954, VAN AKEN and

TASSERON 1955 and 1956, GUTSCHE and SCHROEDER 1963,

MEYNE 1964, TANI 1966 and VAN LAMMEREN et al 1969.

In the figure have also been plotted 20 FPP and 16 CPP-results from open water tests with 11 merchant ship

fourbladed propeller models under cavitating conditions

in the KMW propeller tunnel. The mean value of the

diameter of these propellers is 0,254 m, the smallest

diameter is 0,219 m and the largest 0,267 m. For these

testsCA

is practically independent of the cavitation

number cl'when 2. This is considered in practice to

be the lowest value for merchant ship propellers.

Usually0'Ballast> 3 andFull load > 5. The scattercl

between the results is relatively modest considering that

they originate from several sources. The following

39

Parameters are varied in figure 16:-Z, AE/A0, P/D I/D,

radial pitch distribution, section chamber, rake, hub

diameter ratio, blade thickness and thickness distribu-

tion. From VAN LAMMEREN et al 1969 can be seen that the

CA value is practically independent of the number of

propeller blades.

The relation, between CA and PA/D is Well establiehed

for both FPP and CPP. The difference between FPP and

CPP is explained mainly by the fact that the blades of

a reversed OPP are working, under even more stalled'

conditions than the blades of a reversed FPP, NORDSTROM

1945..

Mean curves from the KMW cavitation tests for FPP and

QPP are shown in figure 16 and these CA values will be

Used Iereafter.

. .

ESTIMATiNG STOPPING QUALITIES WITH THE AID OF REGRESSIONANALYSIS

By putting equations (37) and (38) into logarithmic form

linear relationships are obtained approximately. This

is shown by plotting the stopping results in loglog

diagrams. Analytical expressions have thus been obtained

by using the method of least squares. These relations

can be expressed as

P1 ST

10l10 AOlog

logAs

logAA

Vo

g t1 s 10°log - log At

Vo

where A and B are constants.

Based'on equations (23) and (35) an attempt has been Made

to exchange QA against Q0 also for PPP ST. This simplifies

the procedure When estimating the stopping performance.

PAn10log

QA CA

40

Otherwise the relation nA/no must be known or estimated.

Comparisons show that the accuracy in ST and ts is not

influenced whether the estimation is based on QA or Qo

considering the uncertainty in estimating nA/no. There-

fore hereafter when estimating or checking stopping

performance QA will be exchanged for Q0.

For FPP it is assumed that PA Po. Initially for CPP

PAG/Q0CA was calculated with the actual mean value for

PA from recordings made during the stopping tests.

However, this method is inconvenient as the mean PA

value is not known in advance. Thus it is desirable to

be able to estimate stopping performance for CPP also

when PA is unknown- From the CPP statistical material

the mean value 7 is obtained asA

17- 0,75 PA 0

This implies that the statistical mean value of the astern

pitch during a crash stop is 75 % of the desigh pitch

ahead. Utilizing this relationship about the same accuracy

in the result of the regression analysis has been obtained

as when using actual mean pitch values. Therefore

equation (43) will be used throughout for CPP.

It is now possible to estimate mean vaUes for stopping

performance for ships with FPP as well as CPP Simply

by using values for Po, 6 and Qo. Equations (41) and

(42) will then be slightly modified.-

10 g 10loglog = log as

Vo

g ts10log

10- log at +

0

PAb 10log

Q C0 A

10log

where a and b are constants.

In linear form the expressions are

CA

( 43 )

(44)

(45)

g tsV0

PZ b( ) S

OA

P b+

=at A- )t Q0 CA

The result of the regression analysis for the three

groups FPP DE, FPP ST and CPP DE and ST is shown in

table 2. As mentioned above the statistical material

consists of data from full load and ballast, approach

speed at design value as well as appreciably below.

The correlation factors r are high for the three groups

and Student's t-test shows that the values of the

correlation factors are extremely significant. In

addition the values for the logarithm of one standard

error of estimate S have been given. Based on the r-

and t-values the result in table 2 is considered to be

significant. The exponents bs and bt are highest for

FPP ST and lowest for CPP DE and ST. The steepness

indicates the sensitivity in stopping performance for

changes in PA, L or Q0. The results show that ships

with CPP are the least sensitive to those changes.

From the table it is also seen that the scatter is largest

for FPP ST and smallest for CPP. This result is to be

expected because the material from the stopping tests

indicates an appreciably larger scatter in tR for FPP

than for CPP. The value of tR directly influenoes the

stopping result. The time for changing the direction of

thrust' is closely related to tR. From an examination

of trial results of several ships HEWINS and RUIZ 1954

state that the time to develop maximum astern thrust

TAt is equal to tR. This relation can also be seen

in sJAEGER and JOURDAIN 1968. From the extensive tests

reported by HARA et al 1966 the astern thrust at tR

usually ranges between 0,3 and 0,7 TA. Further in the

discussion of SMITH 1937 Dodson shows from cavitation

tests with a model propellex. at low ahead revolutions

acting as a brake in stopping a vessel that although

- 41 -

42

the propeller cavitates heavily in some conditions the

astern thrust is appreciable. During a crash stop the

astern thrust can reach considerable values when shaft

speed or-propeller pitch still are positive. For a CPP

the astern thrust is high already early during the pitch

reversal independently of the momentary torque reduction

due to the braking effect of the blades. The frequency

distributions of tR are shown in figure 17.

In this connection it is of interest to study the

frequency distribution of the mean time value for nA/no

for the three combinations. Figure 18 gives the distri-

butions. In table 3 the arithmetic mean value and the

Pearson coefficient of variation V for the various

combinations in the figure are shown. The Pearson

coefficient of variation which is expressed in per cent

is one standard deviation times 100 divided by the mean

value. From table 3 it is obvious that nA/no for FPP DE

does not have the high values sometimes stated (5.0 85)

whilenA /n0

for FPP ST is close to 0,5 which is a

wellknown figure. The scatter is slightly smaller for

FPP ST than for FPP DE. With CPP the relative scatter

is less than half the aforementioned values and the

shaft speed astern is close to no. The reason for nA

not being equal to no is that not all CFP's are

reversed at full shaft speed. In 37 % of the tests with

CPP shaftdriven auxiliaries have been in use with full

shaft speed during the crash stop manoeuvre.

Equations (46) and (47) have been drawn in loglog

diagrams, figures 19 and 20, where also the statistical

material has been plotted. The mean lines of the three

groups diverge for increasing PAZ/Q00A for gST/V02 as

well as gts/Vo. This is shown in figure 21 for gST/V02

where equation (46) has been drawn in a linear diagram.

The same tendency is valid for gts/Vo. The curves level

out with increasing PALVQ0CA i.e. for larger vessels.

Based on figures 16, 19 and 20 it can be seen that

- 43 -

superior stopping performance is achieved with ships

equipped with CPP. This is treated below. With a CPP an

appreciably higher stopping energy can be transformed

than with an FPP. With FPP DE the stopping performance

is usually somewhat better than with FPP ST. However,

a strict comparison between FPP DE and FPP ST can only

be made using actual values of PAA/Q0CA. The relation

between these is such that P LS/Q C for DE is equal to,A 0 A

or larger than, the value for ST when comparing ships

of the same size and power.

The scatter of the statistical material in relation to

the respective mean curves has been studied for the

three groups. The y2 test shows that the data approxima-

tely follows the normal curve. The frequency distributions

do not differ significantly from their respective normal

distributions. With the aid of the S value which is the

one standard error of estimate, lines for + S and + 2 Sy

have also been drawn. Between the lines + 5y and + 2 SY

respectively 68 and 95 % of all stopping tests should

fall. This is valid for material strictly following the

normal distribution, see figures 22 to 27. For the

statistic material this is approximately valid, table 4.

The per cent deviation for one standard error of estimate

is shown in table 5.

Diagrams of the type shown in figures 22 to 27 can be

used for appraising stopping test results which usually

are only vaguely commented upon. It is then possible to

estimate on a statistical basis, if the stopping test is

successful or otherwise.

The method of stopping a ship by rudder Cycling has been

mentioned previously-. In figure 23 two crash stops and

the best rudder cycling result with the sate Value of

P Ls/Q C have been plotted for the 193.000 dwt steamA 0 Aturbine. tanker "Esso Bernicia", CLARKE et al 1972. The

44

crash stop results are of medium class and the rudder

cycling result is at the same level as the best crash

stop results for FPP ST. From figures 16, 23 and 24 can

be deduced that the best rudder cycling manoeuvre for

"Esso Bernicia" with 93 % probability will result in a

longer track reach than if the ship is equipped with a

CPP and makes a crash stop. The "Esso Bernicia" best

stopping distance when rudder cycling is 25 % longer

than the mean crash stop value for a corresponding CPP

ship. In figure 22 another previously mentioned compari-

son has been made between crash stop and rudder cycling

with a 150.000 dwt motor tanker. Its crash stop result

is better than the expected dean value and the rudder

cycling is not so good as the mean crash stop value.

COMPARISON OF VARIOUS CALCULATION METHODS FOR ESTIMATINGTI STOPPING PERFORMANCE .

In order to examine the proposed method for estimating

the stopping performance the following'weliknown methods

have been studied; CHASE et al 1957, SAINSBURY 1963,TANI 1968, ILLIES et al 1970 and CLARKE and WELLMAN 1971.

These. methOds assume throughout that the vessel keeps its

course during the stopping manoeuvre. Manifestly the result

should be a longer stopping distance and Stopping time than

the real values. Hereby a safety, factor is said to be .

introduced.

Further it is assumed that the astern thrust is constant

and that it has the same value as the thrust developed

by the screw when backed in open water at the appropriate

torque and at zero speed of advance. The propeller

characteristics are based on model tests in atmospheric

conditions. From trials with the "Esso Suez" and a

number of model tests CHASE et al 1957 use a CA value

of 5,5 for FPP. For CPP in the same reference a CA value

of 4,33 is used. This latter figure is taken from two

- 45 -

full scale observations and three Model tests. When

calculating the stopping performAnce according'to

SAINSBURY 1963 it has been assumed that the TAis Values

from CEASE et al 1957 are valid. CLARKE and.WELLDWN 1971

use KT/KQ2/3 2,25 as an average for tankers. The.

overall propulsive efficiency neCessary to know when

calculating according to ILLIES et al 1970, has been

estimated frOm curves received by letter from

Mentzendorff in1971.. TANI 1968 reduces TA.t.ty 7,5%,while in reality the astern thrust cannot ti realized

itMediately at the stop order. 0LARKE.and WEIJI4AN'

-1971 introduce a thrust deduction of 7 % for the bollard

Pull astern value, based on measurements of thrust in

fUll scale. ILLIES et al 1970 euggests a thrust deduc-

tion coefficient of 15 %.

When calculating according to these methods, except that

of TANI 1968, it has been assumed that the QA/Qo-values

according to CHASE et al 1957 are valid. These are

based on the Troost B 4-40 series with variation in P/D.

The torque limit line of 0,8 Qo at 0,5 no and 1,0 Q0

at 1,0 no is used for FPP ST and 1,0 Q0 for FPP DE

between 0,5 and 1,0 no. For all five methods it is

required to know or estimate the value of nA/no at dead

in the water. The measured nA values for that condition

have been used in the calculation.

The time from the stop order till full astern thrust

has been put at 20 seconds when calculating according

to CHASE et al 1957. This. value has been suggested' in

that paper when actualfigures are not available.

TANI 1968'does not mention the reversing time but

presumably the method takes account of ityorthe.other

three methods the real t values have been Used ih the

calculation.

SAINSBURY 1963 assumes-more correctly that the actual

46

ship speed when the reversal has been effectuated shall

be put into the calculation instead of the approach

speed. This is a complication as the value of the speed

is not known and must be estimated. ILLIES et al 1970suggest a simple formula for estimating the ship speed

at the reversal. The SAINSBURY 1963 method has been

checked with that formula and the result is that the

scatter is appreciably greater than if the approach

speed value is used. Further from the crash stop

recordings it is seen that the actual ship speed when

the propeller is reversed is very close to the approach

speed. Therefore when calculating according to

SAINSBURY 1963 the approach speed has been used. This

is also the case for the remaining four methods.

The ship resistanee is assumed to be proportional to V2

which is approximately valid for the greater part of the

speed range. Ro; i.e. the ship resistance at the stOporder, has been calculated according to SAUNDERS 1957for CHASE et al 1957 and SAINSBURY 1963. TANI 1968 bases

R0 on model tests with tankers above 40.000 dwt and.

CLARKE and WELLMAN 1971 apply a Froude resistance

coefficient . 0,65 as an average for tankers.

According to CHASE et al 1957, SAINSBURY 1963 and CLARK

and WELLMAN 1971 the "added mass" is 8 %. ILLIES et al.

1970 use 10 % of L and TANI 1968 probably uses results

from MOTORA 1960 for "added mass".

ILIJIES et al 1970 utilize the energy relation to solve

the expressions for ST and ts. The Other four methods

use the equation of motion (7). Based on these

assumptions -

9

kx) V:U e RoS = V log. T 0 g

1 +TAt

)2 R0

+ kx

V Rtan-1 V RO

= tRT0 Ats TAts

- 47 -

With the aid of the basic equations (48) and (49) the

special assumptions valid for the various methods have

been applied resulting in different expressions for

STand ts derived by the above authors. For the actual

expressions reference is made to the respective method.

The methods according to CHASE et al 1957, SAINSBURY

1963 and ILLIES et al 1970 make no restrictions

regarding.type or size of ship. On the other hand the

method of OLARKE and WELLMAN 1971 has been detiVed.

especially for tankers and that of TANI 1968 is ,.valid

only for tankers above 40.000 dwt. The, greater part of

the material on which the latter method is based lies

'within the boundaries 40.000 to 70.000 dwt.

The stopping performance for the statistical material

has been calculated by these methods on the above

assumptions. The calculated result has then been

compared With the corresponding real. values. In table

6 a comparison is shown where the results of the above

suggested calculating method are also included. A

graphical 1.epresentation is given in figures 28 to 32.

For FPP tankers and bulk carriers 91 tests have been

carried out in full load. For various groups of tests

the average arithmetic values as well as the relative

variability or scatter expressed by Pearson's,

coefficient of variation are shown for STc/ST and

t /ts The same stopping test material has been usedsc

throughout. The table shows that the degree of accuracy

forSTc /ST

is higher than that for tsc /ts in the five

methods. Further the estimating method suggested here

has approximately the same degree of accuracy regarding

STc/ST as the most accurate of the other methods. When

compared to other methods tsc/ts has significantly less

scatter. In addition the calculation for vessels with

CPP is more accurate than for those with FPP. Three of

the five methods indicate shorter-stopping distances

than those recorded although no resistance increase

due,to turning is taken into account in any of the methods.

- 48 -

Another comparison 'between the various methods is shown

131 figures 33 to 37. There the material, has been divided

into two groups namely

g >1, Vo 13 knots

a() < 1 or Vo < 13 knots or af < 1 and Vo < 13 knots

From the above it is seen that the sheer of the vessels

is significantly less for the latter than for the

former group. According to figures 33 to 36, for

vessels which do not sheer appreciably, the five methods

give estimates of track reach and stopping time which

in approximately 60 to 80 % and 60 to 70 % respectively

are lower than the average values for the combination

"all ships" in table 6. The corresponding figure based

on the method suggested above is approximately 50 %.

The reason for the calculated values being lower than

those measured can mainly be attributed to that the

propeller thrust in reality is influenced by cavitation

and air drawing not taken into account in the calcula-

tion. For the cathegory > 1, Vo? 13 knots on the

other hand the majority of the estimated values fall on

the high side of the mean value for "all ships". This

is due to the turning resistance not being considered,

the influence of which can be large. Again for the

method introduced here approximately the same number

of estimated values are on both sides of the mean value.

The influence of the resistance increase from the turning

motion as well as the time for reversal, tR, is inherent

in the method suggested here. Further its application

is relatively simple as only the following factors

require to be known or estimated:

approach speed Vo

propeller design pitch ahead Po

propeller diameter D

ship displacement

propulsive machinery design torque Q0

- 49 -

For CPP, Q0 is to be corrected for shaft driven

auxiliaries if any.

TEE STOPPING EQUATION FOR ONE PER CENT RISK

The principle used based on statistics makes it possible

to calculate a curve for which the crash stop distance

is not exceeded in more than for instance 1 % of all

cases. This is considered to be a reasonable figure.

In other words there is a 99 % probability that the

vessel will stop in a shorter distance. Based on the

normal distribution this corresponds to a curve whichis 2,33 S above the mean curve. These so called risk

curves which in the log-log diagrams are parallell to

the mean lines, will have the as and at values shown in

table 7. The bs and bt values are identical to those in

table 2. In figures 22 to 27, the risk line and a line

which indicates that 1 % of all crash stops will be

shorter, have been drawn.

In practice, it can also be of interest to know the

1 % risk value for SH. Such risk lines for head reach

have been calculated, using results from tests where

g < 1. As mentioned above the sheer for e < 1 is

appreciably less than for g >1. This implies that the

SH/ST values are significantly closer to unity fore < 1.

Aboard it is not a simple task to determine if o > 1

or < 1 and consequently when estimating a risk distance

along the initial course, SH for g< 1 has been chosen.

The result of the calculation is that the 1 % risk

SH-line is close to the 1 % risk S -line but it is not

so accurate. Studies of the frequency distribution from

plotted values of SH/ST above 14 knots, show-that there

is close agreement between the groups PPP DE, FPP ST and

CPP DE and ST. Below 14 knots the Si/ST values are

dependent on approach speed and are appreciably higher

because at reduced approach speed the sheer is much lower.

The material from the three groups has been put into two

sL -25V

- 50-

frequency distributions, namely with g > 1 and e < 1the characters of which are distinctly different, see

figure 38. The median value of SH/ST in the figure is

approximately 0,68 and 0,94 for 0 > 1 and <1 respec-

tively. Thus for g < 1 50 % of all SH-values are within

6 % of ST. Therefore and due to the larger scatter in

the SH-values, it is considered preferable to use the

STrisk line in practice when estimating the risk

distance ahead.

In the same way as for ST and SH, attempts have been mad

to calculate the equations for gSL/V02 as a function of

PAA/Q00A. This, however, resulted in a too large1/3

scatter. Instead the quotient STIP has been studied.

From figure 4 it is seen that the highest Slip1/3

values are reached for Qtf > 1. Since the purpose is to

determine the lateral risk distance, the material for

g > 1 has been selected. At full load and with an1/3

approach speed close to the design value, SL/ is

not dependent upon size. Frequency distributions for11/v do not show any characteristic difference

between FPP DE and ST or between SB or PT sheer. There-

fore it has been considered acceptable to treat FPP DE

and ST, SB and PT sheer as one material. There is a

dependence on approach speed up to approximately 13

knots which is seen in figure 39 where SL/s71/3 is

plotted as a function of Vo for FPP DE and ST, SB and

PT, g > 1. Above 13 knots there seems to be no speed

influence. The frequency distribution above 13 knots

has been studied. Its mean value is expressed as

(V-0 13 knots) (50)

and the coefficient of variation is 45 %. The value

SL- 51 (Ir 13 knots)

will not be exceeded with more than 1 % probability.

( 51 )

5 1

For Vo < 13 knots it has been assumed that

"%0VO

Then the following is obtained:

SL

773 0,30 V20-

This relation is suggested to give the lateral dis-

placement for 1 % risk.

FQT CPP On the basis of the same line of argument

the mean value is

sL

7 (v0 13 knots)7177 -4

The coefficient of variation is 32 % and the 1 %

risk value is

SL773 - 30V

( < 13 knots)

(1/.0 2_13 knots)

( 5 2 )

( 5 3 )

( 5 4 )

( 5 5 )

In the statistical material there are no stopping

tests with CPP and 4P> 1 for V0 K'13, knots. Using the

SaMe reasoning as above for FPP; the following relation

is valid for CPP at approach speeds below 13 knots.:

V1 / 3 '18 V2 (56)

The difference In SL/ mean values between FPP and

CPP is significant.

PRACTICAL CRASH STOP DIAGRAMS

The expressions for 1 % risk can be used on board so

that the vessel can be handled with small risk of

collision._Such diagrams, approximately valid in full

load as well as in ballast, are shown in figures 40 to

52 --

43 for two sisterships, bulk carriers of 75.000 dwt,

one with FPP and one with CPP. For the ship with CPP

the power at the propeller is 3 % lower than for the

ship with FPP as in the former ship a shaft driven

generator is used. For instance at full load, 95.000

tons and 16 knots approach speed the FPP ship in 99 %

of all crash stops will have a stop track of less than

2,7 nautical miles. The corresponding distance for the

CPP ship is 1,6 miles. Thus there is more than 1 mile

of extra margin for the CPP ship. For lateral deviation

with the CPP ship the 1 % risk distance is 0,7 miles and

for the FPP ship 1,2 miles.

It is suggested that the SVCsdiagrams shown in figures

40 to 43 be used for instance as follows for the two

ship case, see figure 44: The master of ship A deter-

mines the distance AB to ship B. Master A shall assume

that ship B will not stop in a shorter distance than

ship A. Master B is to reason corresPondingly that ship

A will not stop in a shorter distance than ship B.

Suppose for example that A and B are the sisterships

mentioned above and that. ship A has OPP and ship B FPP.

Further. assume that both are fully loaded. Master A

whose ship, at the radar contact with ship 33 makes 12

knots, can keep this speed as long as AB:>2 STmin = 1,8:

When,AB closes up to 1,8'master A can order a crash stop.

Ship B which makes 12 knots at the radar contact is

manoeuvred correspondingly when AS 3,0'.'

On board it is considered important-to be able to

estimate the stopping time. The longer ts is, the longer

it will be impossible to manoeuvre the ship. The slip

cannot be steered during the stopping manoeuvre and it

is natural to try to minimize the period the vessel is in

this condition. Figures 45 and 46 show practical tWs-

_diagrams for the stopping time of the mentioned 75.000

dwt ships. In 99 cases out of 100 the ship with OPP is

- 53 -

expected to craSh. stop from 12 knots fully loaded in

less than 8 min 45 sec. For the FIT ship the corresponding

time at risk is 5 minutes longer.

Earlier it was mentioned that Oosterveld has developed a

practical stopping diagram, figure 1. Diagrams of this

type have been distributed to various classes of Shell

tankers and are said to give representative mean values.

From SHELL 1968 the stopping results from design speed

and full astern have been plotted in figures 22 and 23.These show that the crash stop values are on the mean

lines for the respective groups. This implies that in

50 % of all cases the crash stop result can be longer

than that indicated in the data for the Shell tankers.

It will be appreciated therefore that this type of

diagram does not give the master sufficient information

on the distance and time at risk.

COMPARISON OF STOPPING PERFORMANCE WITH FPP AND CPI'

SHIPS

Amongst those who have compared the stopping performance

-between.ships:with FPP and CPP are HOOFT and VAN MANEN

1968, RITTERHOFF 1970, MASUYAMA 1970 and 1971 and OKAMOTO

et a1 1971. In table .8 results from those studies have

been collated. The results vary appreoiably.

The study by HOOFT and VAN MANEN 1968 is based on

stationary model tests with reversed propeller. In

addition in that investigation, the full scale propeller

design shaft speed for ST as well as DE has been put at

85 revs/min. Further the diameter for the FPP is 9,193 m

and 9,480 m for the CPP. The comparison is not considered

to be fully realistic but not withstanding this it has

been referred to and treated by several authors, among

others by RITTERHOFF 1970 who interprets the results

somewhat differently, see table 8. The latter reference

for crash stop (quick reversal) gives relatively high

ts CPP1,245

ts FPP

- 54 -

figures of ST app/ST Fpp and ts cpp/ts Fpp for DE, as

do HOOFT and VAN MANEN 1968. MASUYAMA 1970 bases the

figures on full scale observations as well as theore-

tical calculations but in the 1971 reference uses only

full scale results. The statement by OKAMOTO et al 1971

is founded on full scale crash stop tests with two high

speed CPP DE cargo liners and a CPP ST OBO-carrier of

130.000 dwt. In this latter reference it is claimed that

a CPP only improves the stopping qualities for ships

with DE and not for those with ST: "The difference of

stopping performance between FPP and CPP is expected to

be small". This statement is not in accordance with

observations from ship trials.

With the aid of the mean curves based on the statistical

material, equations (44) and (45) the relations

ST CPP/sT FPP and ts C/tPPs FPP have been studied. It

has been assumed that the following factors are

reciprocally equal: P0, A , Q0, no and D. Further the

relation -FA = 0,75 Po has been used. The material covers

the size range from 570 to 134.000 dwt. Actual data for

ship, machinery and propeller have been utilized for

full load and ballast and the result has been divided

into DE and ST groups. A suitable way to present the

above relations is to use (PAL/Q OCA)FPP as a basis.'

For the vessels studied the mean value of (PAL/Q0CA)cpp/

(PAZN/Q0CA)Fpp is 1,33 for DE and 1,31 for ST. With these

figures and equations (44) and (45) ST app/ST Fpp as

well as ts cpp/ts Fpp can approximately be calculated

as a function of (PALI/Q0CA)Fpp. Thus the following

indicative curves are arrived at

go C 0,068 'sST CPP_ 1,045 ( (57)

sT FPP A FPPDE

0,088

FPP(58)

s FPP

These equations are presented in log-log diagrams,

figures 47 and 48 together with plottings for ships in

full load to indicate the influence of size. The

numbers at those points represent dwt/1.000. Due to the

divergence of the mean curves the difference in stopping

performance increases with increasing PAL/Q0CA, whioh

generally means increasing ship size. For the statistical

material, up to approximately 130.000 dwt, the mean data

in table 8 is valid.

For the lateral displacement also a comparison can, be

'made between TPP and CPP ships. From above the mean of

SL/ \-71/3 is 25 and 17 for FPP and CPP respectively.

Thus SL cpp/SL Fpp"--= 0,68 as a mean value for e > 1.

SUMMARY

- 55-

An attempt has been made to solve the stopping distance

and time with the aid of a large number of observations

from crash stop tests. The material studied consists of

330 tests from single screw vessels built at 25 shipyards

in 8 countries. 75 of the tests are from ships with CPP.

The size and type range from dry cargo vessels of 600

dwt to tankers of 230.000 dwt in ballast as well as full

load, figure 2. The statements made refer to the

material used in this study.

Based on the criterion for dynamic stability at straight

line propulsion ahead. by WAGNER SMITT 1970 and 1971, an

approximate stability criterion for vessels at crash

stop has been developed, figures 4 to 8 and equation (6).

ST CPP (

(

c. 01 ,54)

> ST

(59)

(60)

1,592T FPP

ts CPP

pu

A FPP

Q,0 cA 0,154- 1,493 )FPP

56

This criterion indicates whether the sheer Of the vessel

will be moderate or whether there is risk of large

deviation from the Original course. The criterion is

valid for trimmed as well as untrimmed vessels. It some

parts of the study the Criterion is used to divide the

material into stable and Unstable vessels.

The probability of sheer SB and PT is given for three

categories, namely vessels with FPP DE, FPP ST and CPP

DE ST, figures 9 to 11. The influence of dynamic

stability, direction of wind, propeller type and

machinery has been considered.

The turning angle at crash stop is shoWt for stable and.

unstable vessels with FPP or:CPP, figure 12. The diffe-

rence in mean value of turning angle between Ve'Ssels

with.FPP and OPP is not significant.

The influence on retardation during crash stop from

ordinary ship hull resistance, turning resistance, wind

resistance, propeller astern thrust and uncertain

factors is discussed. It is shown that for practical

estimation of crash stop performance it is sufficient

to put ST and ts as functions of Vo, PA,A , Q0 and

CA.The astern thrust constant CA is given for FPP and

CPP as a function of PA/D from open water tests as well

as from KMW cavitation tests, figure 16. It is suggested

that the curV'es representing cavitating conditions be

used.

The stopping qualities have been studied with linear

regression analysis for the groups FPP DE, FPP ST and

CPP DE ST. The relations have been put in logarithmic

fort and it is seen that there is approximate linearity

for the logarithmic expressions, equations (44) to (47),

table 2, figures 19 and 20. The scatter of the statistical

material in relation to the respective mean curves has

been studied for the three groups, tables 4 and 5. The

57

2'X test shows that the data approximately follows the

normal distribution. Curves for +1, 2 and 2,33 standard

errors of estimate (Sy) have been evaluated. Material

for estimating and appraising crash stop performance for

FPP and CPP has thus been obtained, figures 22 to 27.In these diagrams results from rudder cycling tests

can be compared with results from crash stop tests.

The curve with + 2,33 S shows that the crash stop

result can be longer than the diagram value with 1 %

probability. Practical diagrams for each ship can thus

be designed based on the 1 % risk curve, equations (46)

and (47), table 7, figures 40, 41, 45 and 46.

The expressions on ST and ts from equations (46) and

(47) have been compared with measured values from the

crash stop tests. This has also been done for five well

known methods, namely those by CHASE et al 1957,

SAINSBURY 1963, TANI 1968, ILLIES et al 1970 and CLARKE

and WELLMAN 1971. Frequency diagrams of STc/ST and

tsc/ts are shown, figures 28 to 37. Mean values of

STc/ST and tsc /tshave been calculated as well as the

relative variation, table 6. It is seen that for the

material investigated (156 tests) the estimating method

developed in this study is the most accurate. This is so

although for the five methods with which the comparison

was made, the measured nA values at dead in the water

and for three of the methods the measured tR values

were used in the calculation.

Head reach, SH and lateral reach, SL have been discussed.

For safety reasons it is suggested that the track reach

values ST be used as head reach. Expressions for

S /pi are given for FPP and CPP vessels', equations

(50), (51), (53), (54), (55) and (56). Individual

diagrams for lateral reach for each ship, based on the

1 % risk value, are shown, figures 42 and 43.

58 -

Comparison of crash Stop performance is made between

vessels equipped with CPP on the one hand and FPP DE

as well as FPP ST on the other, based on mean value8,

equations (46) and (47) and table 2. Relations for the

comparison are given, equations (57) to (60). The crash

stop qualities for vessels with CPP are clearly superior

to those for vessels with FPP. The reason for this is

that with a CPP the thrust reversal is quick and the

backing energy is higher than with an FPP. The

difference becomes larger with increasing ship size,

figures 47 and 48.

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HORNE, L R:

- 61 !-

HARVALD, S A: "Wake and thrust deduction at extremepropeller loadings", SSPA Publication Nr 61, 1967

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ILLIES, K,, =ERSE, G, MENZENDORFF, H and RITTERHOFF, J:"Analyse des Verhaltens Von Schiff und Maschine bei -

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large1968

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KEMPF, G and HELM, .:. "Auslaufmessungen am Schiff undam Modell DaMPfer:'HaMbUrge", WRH.,'1928_

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- 63-

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TABLE 1

Data for evaluating the stopping_performance ofns

Yard number /23Ship type .7.--0M/KgR

LOA(length over all) 26441M

L between perpendiculars)

4,1B (moulded beam)PA

M d (mean draught) ./.6/ 6 f IIdwt (tons dead weight) ./.Aq .Q9.Q..V (volume displacement at mean draught)

.24/6OOOOOOOOOOO

cf3

P Type of machinery (steam turbine or diesel engine) .4.99 SHP (shaft horse power, design. For CPP ship also design

0 power of shaft driven auxiliaries) t211 009

Type of propeller (FPP or CPP) F PPRight- or lefthanded propeller R/6.HTZ (number of blades)

di

D (propeller diameter) 1.4:679./. O

k P0 pitch at 0.7 radius)0 uH

AE /A0(expanded area ratio) .P4444;4

T-1 /7a RPM° (propeller revolutions per minute, design) .... OOOO

0P V (speed of ship, design) Adir. c?!".crnsa D

E (distance from propeller shaft centre to ship's base

line) A.A-47.4?.!1. OOOOO

dF (draught at forward perpendicular) 614416.1 OO .

dA (draught at aft perpendiCUla±)

.(weight displacement) 9.4??. .........

Wind speed and direction . .....

Sea .....

Waterdepth 170

p., Vo (approach ship speed at stop order

0 t (stopping time) acsi o'rpti

-P sV = fi (t) and RPM f2 (t)

Ship's course during stopping

TABLE 2

Result of regression analysis of equations (46) and (47) for mean track reach and stopping time

NaS

bS

rS

tS

10g

lo

s Y S

at

bt

rt

1°log

Sy t

FPP, DE

82

4,44

0,684

0,93

22

0,094

7,80

0,693

0,93

23

0,090

FPP, ST

79

2,88

0,770

0,86

15

0,100

6,44

0,759

0,89

17

0,086

CPP, DE and ST

46

3,89

0,616

0,97

25

0,066

8,17

0,605

0,97

25

0,067

TABLE 3

Mean value and variation of reversed shaft speed

Pearson'scoefficient ofvariation %

TABLE 4

Distribution in % of the total number of observations .

within the limits + 1 and + 2 standard errors of estimate

respectively

gST/v20 gt vo

+ S +.2 S + S + 2 S_ _ _

TABLE 5

Variability in % for one standard error of estimate

g 5T1vo2g t5/v0

FPP, DE 0,65, 16

FPP ST 0,53 13

CPP, DE and ST 0,96 5

PPP, DE 67 95 .66 95

FPP, ST 63 96 68 96

CPP, DE and ST 70 96 67 93

FPP, DE 22 21

FPP, ST 23 20

CPP, DE and ST 15 15

TABTR 6

Relation between calculated

and real values of stopping

distance and time

Number

Combination

of tests

8Tc/

8TPearson's

Average

coefficient

value

of variation %

tB

Cit

Average

value

BPearson's

coefficient

of variation

PPP,

all Ships

156.

0,76

825

,80,

882

30,9

FPP, tank + bulk

113

0 81

624

,50,

959

28,6

CPP

, all

ship

s29

0,91

822

,31,

034

26,0

PP, tank + bulk

250,

941

21,8

1,05

526

,2

FPP, all ships

-

156

0,94

825

,30,

925

28 4

-

FPP, tank + bulk

11.3

0,94

724

,10,

963

28,5

FPp, all ships

156

0,93

138

,51,

188

49,2

FPP, lank + bulk

113

1,02

834,0

1,31

345

,8

PPP, tank + bulk

95

1,04

035

,51,

341-

47,9

>40.000 dwt

FPP, all ships

156

0,74

437

,80,

964

39,8

FPP, tank + bulk

113

0,81

932

,71,

075

34,3

FPP, all Ships

156

1,09

025

,01,

167

31,4

YPP

1tank + bulk

113

1,12

0.23

,61,

249

30,1

PPP,

all

ship

s15

,6,

0,99

523

11,

026

20,0

FPP, tank + bulk

113

0,99

9:2

3,1

1,01

520

,6

CPp, All

ship

s46

1,01

115

,31,

007

15,5

OPP, tank+ bulk

321,

003

17,4

0,99

416

,6

TABLE 7

The

1 % risk curve coefficients for track reach

and stopping time in equations

(46)

and

(47)

bt

bS

0,69

3

0,75

9

- 0,

605

FPP,

DE

7,35

0,68

412

,65

FPP,

ST

4,93

0,77

010

,21

CPP, DE and ST

5,55

0,61

611

,70

1>crash stop

2?

up to r-4130.000

dwt

TABLE 8

Comparison of estimates of stopping performance between PPP and OFF ships

Reference

st

CPP

ts CPP

Notations

ST

FPP-

tsPP

HOOFT and

VA

N M

AN

EN

1968

0,69

0,85

ST, 100.000 dwt

0,94

1,18

DE, 1[00.000 dwt

.MASUYAMA

1970

,a0,70

^,0,70

ST, 120.000 - 130-.000 dwt

- 0,70

DEand ST > 20.000 dwt

RIT

TE

RH

OFF

1970

0,70

0,83

ST, 100.000 dwt

0,58

0,86

DE

,10

0.00

0 dw

t, sl

ow r

ever

sal

0,87

1,13

DE

, 100

.000

dw

t, qu

ick

reversal

MASUYAMA 1971

r-d0,65 -

0,70

r-0,65 - 0,70

ST, 120.000 - 130.000 dwt

OKAMOTO et al 1971

1.

1ST, 130.000 dwt

0,5

0,5

DE

,6.000 - 14,000 dwt

NO

RR

BY

1972

0,55 -

0,71

0,52 - 0,67

ST, 1000>P0G/Q0C0150

2)

2)

0,65 - 0,75

0,68 - 0,82

DE, 1200>P00 CA > 95

'

2,0 0

30

,;c\

-c\..

r`f)

1012

141

6085

Figu

re1. Oosterveld

diag

ram

for

determining stopping

dist

ance

And

sto

ppin

g tim

e.

18V

0 kn

ots

RPM

60

2

No of tests100

11CPP, Steam turbine,

FPP, Diesel engine,and bulk carriersFPP, Steam turbine,and bulk carriers

4 tests with OBOcarriers

124 tests, 65 with tankers

131 tests, 126 with tankers

.3dwt x 10

20 40 60 80 100 120 140 160 180 200 220_240.

Figure 2. Stopping tests, frequency of distribution.

CPP, Diesel engine,and bulk carriers

71 tests, 34 with tankersI

I

IM

Full astern

Figure 3. Definition of head reach' SH and lateral

reach, SL'

0

AP

50 Figu

re 4

.

0 0*

Acf

r A

ktL

P 04

.4A

A *

L.

a

0

(V/d

3 )/

(V/d

3A)c

1,6

0 0

0 0

1,4

0 A0

Ic)

00

0it

A,

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as a

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for

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ps(

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DE

, ZS

: PPP

ST

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l: O

PP D

E S

T (

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ght h

and

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R

30 B

B40

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nd P

T.

turn

ing

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elle

r).

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AA

OC

)0

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AA

®A

I 2'

A A

A i

A

A

I_ III!! III

II

220 200 180 160 140 120 100 80 60

40

20

PT

( V/d A)/( V/d

)

6 oA0 0

7-1

°

0

o0A

e

o0

0.4t

D00

rc' A

_

opo,

o

0 21 20

40

00

1111111 II

60

80 100 120 140 160 180 200 220

Lli°

SB

Figu

re 5

- T

urni

ngangle. T

asa

func

tion

of(S7/d3A)I( S7/(13'.1

i)c

for

ship

s tu

rnin

g to

SB

and PT

():

FPP

DE

,A :

FPP

ST,A :

CPP

DE

ST

(L

left

, R r

ight

han

dturning propeller),

0 t

0-

00

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A01

2A

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3A)/

(V/d

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1 , 6

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0,

A0

A 0

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00

ra09

° °

°di

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4 A

AA

a,ia

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4lb

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1a

it c)

0 °

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0 A

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AA

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1,00

0

AA

1I

II

[0,

2I

5040

3020

100

10F

0 AA

AA

A p

0A

0

A

Figu

re 6

, SL

/ V1/

3as

a f

unct

ion

of (

S7/

d3A

)/(N

7/d3

A)c

for

ehi

ps, l

uffi

ng a

nd f

allin

g of

f fr

ozn

win

d.fl

: FPP

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T. d

i A:

PP D

R

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QD

C9

---

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11

L

203.

040

50

0.1

00

AA

Aac

o °

A

Figure 7. Turning angle y as

function of (V/d

AY

( V

/d3A

)cfor ships luffing and falling

Off fram wind.

0:FPP DE,

LS:

FPF

AL

:CPP DE ST_.

( V/d

1,6

)/(N7/d3A)C

1,4

lo

000

00

A

0 A0

0 go

d,0

0 0

cg

0AA

A

A

A1,

2A

AA

A A

C.0

40A

Atill

A4

ASi

0 A

dr)

: A

0A

00

0

000

0 0

0

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0A

0 40,

0

00

044?

0000

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Opo

4

A

11

10,21

11

11

11,

11

k1

60

40

20

20

40

60

80

100 120 140 160 180 200 220

If°

11

11

11

11

220 200 180 160 140 120 100 80

Figure 8. Criterion formanoeuvres. Stability criterion: Wd3A<( V/d3A)cA

25 50 7 5 100 125 150

V/d3A)CA

FPP_ ST

FPP DE

CPP DE ST

2fe >

44% 56%

37% 63%

30%

22% 7_8 %

70%

Figure 9. Probability of sheer to SB or PT during a crashstop, taking into consideration type of propeller,machinery and propeller rotation effect .(circled.figures indicate number of cases investigated).Assumption: Propeller is left hand rotating afterreversal.

32% 68% 18% 82%

FPP ST

FPP DE

CPP DE ST

26% 74%

wind

24% 76%

20%

50%

wind wind

60% 40%

50%

35% 65%

wind

Figure 10. Probability of sheer to SB or PT during a crash stopfor be> 1 taking into consideration type of propeller,machinery, propeller rotation effect and directionof wind (circled figures indicate number of cases inves-tigated). Assumption: Propeller is left hand rotatingafter reversal.

FPP ST

FPP DE

CPP DE ST

25% 75%

14%

wind

86%

43% 57%

wind

<

60%

70%

92%

Figure 11. Probability of sheer to SB or PT during a crash stopfor 4t4.1 taking into consideration type of propeller,machinery, propeller rotation effect and direction ofwind (circled figures indicate number of cases investi-gated). Assumption: Propeller is left hand rotatingafter reversal.

PT 041N = 13

T's = 24°

SB

N = 44

T= 53°

PT

N=3

af<1

Figure 12. Turning angle

Voat 13 knots

-

40-values are

FPP

CPP

.PT eb1N = 33

T.= 112°j

PT

N= 5

4' for FPP- and CPP-ships,

Approximately 67 % of all

in the shaded sectors.

0

0 ts

Figure 13. Relation between ship speed and time duringa crash stop. Track reach ST is representedby shaded area below 1,7"-t curve.

11-1. -

centrifugal force

Figure 14. Derivation of the longitudinal component ofthe centrifugal force on a turning ship.

awmean/amean

0,20

0,15

0,10

0,05 0

95%

05

ia15

Vwind speed mjsec

Figure 15. Influence of wind resistance on total retardation during crash

stop

.Wind on bow, no turning, Vn = 16 knots. 67 and 95 % of all wind speeds

observed are below 7 and 1

m/sec respectively.

CA

1111

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0,4

0,5

0,6

0,7

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Figure 16. CA as a function of PA/D at bollard pull astern free running testswith models of

FPP and CPP under atmospheric

((A

Pand cavitating

(AA

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Mean lines for FPP and CPP based

nnYmw

% N

(to

tal o

bs)

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DE

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= 8

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030

6090

120

150

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ec

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= 7

4.

030

6090

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150

180

210

240

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ec

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re 1

7. F

requ

ency

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trib

utio

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tR f

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PP D

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PP S

T a

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PP D

E S

T.

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= 4

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ec

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0,7

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0,3 0,4 0,5

0,6

0,7

0,8

0,9

1,0

1,1

nA/n0

nA/n0

nA/n0

Figure 18. Frequency distribution of nA/no for FPP DE,

FPP ST and CPP DE ST.

FPP ST

N = 72

.CPP DE ST

N = 43

% N (total obs)

50.

40

FPP DE

N = 84

30 20

10.

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/V0

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emum

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ni u

nlim

mul

l lio

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num

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1MIM

IUM

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IMM

IUM

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OM

BIl

ltE

NE

NM

I M

ID 1

1111

IM

I M

C M

tn M

IIIM

111

111

HU

I H

IM M

O N

M O

M M

I M

I II

I 11

111

111

II N

IM

E

100

150

200

300

400

500

1000

1500

200

0-1

7A

Figu

re 2

7.gt

s/V

o as

a f

unct

ion

of 7

5AL

/Q. 0

CA

, CPP

DE

ST

, N =

46.

17A

= 0

,75

Po(I

rI0,

605

Mea

n cu

rve:

gt5

/V0

= 8

,17

AL

/Qoo

d

20

15

10

5

0

20

5

0

0 0,2 0,4 0,6 0,8 1,0 1,2 1,4 1,6 1,8 2,0

sTelsT

I I

CHASE et al 1957

-II_ I I ri I

SAINSBURY 1963

AI ill I I In I

0 0,2 0,4 0,6 0,8 1,0 1,2 1,4 1,6 1

STc/ST

TANI 1968_

8 2,0 2,2 2,4

0 0,2 0,40,6 0,8 1,0 1,2 1,4 1,6 1,8 2,0 2,2 2,4-2,6

ST/ST

Figure 28. Frequency distribution Of STc/S2 according:to

various methods of calculation for ships wfth

FPP.E1.77- tankers and bulk carriers, N 113,

AIN: general cargo liners, N = 43.

30

25

20

15

1 0

5

0

15

10

15

10

20

15

10

15

10

'I

0 0,2 0,4 0,6 0,8 1,0 1,2 1,4 1,6STc/SE

Figure 29. Frequency distribution of STc/S.T according to

various Methods of calcUlatiOn for ships with

FPP.E=1: tankers and bun( carriers, N. = 113

general cargo liners, N,= 43

CLARKE and WELLMAN 1971

I I i Fil Mr -1-61 l iA 1 ri

0 0,2 0,4 0,6 0,8 1,0 1,2 1,4 1,6 1,8 2,0 2,2 2,4 2,6

STc/ST

n.

ILLIES et al 1970

0 0,2 0,4 0,6 0,8 1,01,2 1,4 1,6 1,8 2,0 2,2 2,4

ST/ ST

20

NORRBY 1972

20

15

10

20

15

10

15

10

5

n n i i i_ m0,2 0,4 0,6 0,8 1,0 1,2 1,4 1,6 1,8 2,0 2,2 2,4 2,6 2,8 3,0

tsc/ts

mAs0 0,2 0,4 0,6 0,8 1,0 1,2 1,4 1,6 1,8 2,0 2,2 2,4 2,6 2,8 3,03,2

t /tSc s

0 0,2 0,4 0,6 0,8 1,0 1,2 1,4 .1,6 1,8 2,0 2,2 2,4 2,6 2,83,0 5,4 5

tsc/ts

Figure 30. Frequency distribution of t5/t5 aCOording,to

tiethOds of calculation for Ships with,FIT,C=D: tanketSand bulk. carriers, N = 113,OW,gerieral cargo linets'N = 43,

CHASE et al 1957

SAINSBURY 1963

;TANI 1968

15

10

5

0

0 0,2 0,4 0,6 0,8 1,0 1,2 1,4 1,6 1,8 2,0 2,2 4,0 4,2

tsc/ts

0 0,2 0 ,4 0,6 0,8 1,0 1,2 1,4 1,6 1,8t /tsc s

Figure 31. Frequency distribution of t5/t5 according to

various methods of calculation for ships with

FPP 1=3: tankers and bulk carriers, N = 113,

MOB: general cargo liners, N. = 43

CLARKE and WELLMAN 1971

ILLIES et al 1970

0 0,2 0,4 0,6 0,8 1,0 1,2 1,41,61,8 2,0 2,2 3,0 3,2

tsc/ts

25

20

15

10

5

0

15 -

10

10 5 oL

I".

n--1

rr

CHASE et al 1957

N =29 (i=3:

25,111W 4)

00,2 0,4 0,6 0,8 1,0 1,2 1,4

00,20,4 0,6 0,8 1,0 1,2 1,41,6 1,8

10 5 0

-n m

inm

II

1111

n1

0 0,2 0,4 0,6 0,8 1,0 1,2 1,4 1,6

00,2 0,4 0,60,8 1,0 1,2 1,4 1,6

STo/ST

tsc/ts

Figure 32. Frequency distribution of STe/ST and tsc/ts according to CHASE et al 1957 and

NORRBY 1972 for ships with CPP.1:21: tankers and bulk carriers

OW

general cargo

liners.STcl

TtSCItS

NORRBY 1972

N46 q=2t 32,:14)

flfl

20

15

10

0 I El IP

0 0,2 0,4 0,6 008 1,0 1,2 1,4 1,6 1,8 2,0 2,2 2,4 2,6

STcST

Figure 33. Frequency distribution of STc/ST according to

various methods of calculation for ships with

FPP.M: cX> 1, 13 knots, N = 75,1111: aeL.1%

Vo 4 13 knotq/A < 1, Vo< 13 knots, N = 81.

I II n

0 0 20,4 06 0,8 1,0 1,2 1,4 1,6 1,8 2,0

ST/ST

CHASE et al 1957

SAINSBURY 1963

iL I

0,20,4 0,6 0,8 1,0 1,2 1,4 1,6 1,8 2,0 2,2 2,4

STc/ST

30

25

20

15

10

5

20

15

10

5

0

10

5

4 we--

CLARKE and WELLMAN 1971

1 I 1 II 161 1 1 l ri

0 0,2 0,4 0,6 0,8 1,0 1,21,4 1,6 1,8 20 2,2 2,4 2!,6

S /STc' T

ILLIES et al 1970

11- n n 1 1I n I.10,2 0,4 0,6 0,8 1,0 1,2 1,4 1,6 1,8 2,0 2,2 2,4

STc/ST

NORRBY 1972

0,2 0,4 0,6 0,8 1,0 1,2 1,4 1,6

S /STc T

Figure 34. Frequency distribution of STc/ST according to

various methods of' calculation for ships with FPP.

1=3: et> 1 , V1 3 knots, N = 75, IEN: 041/VO413

'1nots/04. 1, Vo < 13 knots, N = 81.

15

10

5

20

15

10

5

0

20

15

n

20

15

10

5

10

5

I I

CHASE et al 1957

TTT1 ni 1 1 1 1 In

0 0,2 0,4 0,6 0,8 1,0 1,2 1,4 1,6 1,8 2,0 2,2 2,4 2,6 2,8 3,0

tsc/ts

SAINSBURY 1963

n I I I I 1 I Int

0 0,2 0,4 0,6 0,8 1,0 1,2 1,4 1,6_1,8 2,0.2,2 2,4 2,6 2,8 3,0 3,2

t /tSc s

TANI 1968

0,2 0,4 0,6 0,8 1,0 1,2 1,4 1,6 1,8 2,0 2,2 2,42,6 2,83 5,4..5,6

ts/tc s

'Figure 35. Frequency distribution of tsc/ts according to various methods

of calculation for ships with FPP.C=D; 4)1.1, \To>, 13 knots,

N = 75,11111: R4:. 1/VO4(.13 knots/441, Volt: 13 knots, N = 81.

20

15

10

5

20

CLARKE and WELLMAN 1971

15

10

0

15

10

5

25-

20

15

10

5

0,2 0,4 0 6 0,8 1,0 1,2 1,4 1,6 1,8 2,0 2, 4,0 4,2tsc/ts

0,2 0,4 0,6 0,8 1,01,2 1,4 1,6

t tsc s

ILLIES et al 1970

NORRBY 1972

0 L I I rri0 0,2 0,4 0,6 0,8 1,0 1,2 1,4 1,6 1,8

tsc/ts

Figure 36. Frequency distribution of t5/t5 according to

various methods of calculation for ships with

FPP.r-1: 0> 1, Vo al: 13 knots, N = 75,.?/(. < 1/V0 < 13 knots/IA < 1, Vo .4. 13 knots, = 81.

11-1

fin

00,2 0,4 0,6

0,8

1,0

12 1,4

STc/ST

Figure 37

CHASE et al 1957

N = 29 (ED:

17,111t 12)

NORRBY 1972

N = 46 (1=1:

22,11111:

24)

nn

10

110

5[

5E

01

i1

n1

01ili

a4

n1

0,2 0,4 0,6 0,8 1,0 1,2 1,4 1,6

00,2 0,4 0,6 0,8 1,0 1,2 1,4 1,6

tSic/ST

efts

0I

IW

1111

11-1

1ri-

rn171

I

00,2 0,4 0,6 0,8

1,0 1,2 1,4 1,6 1,8

tsc/ts

Frequency distribution of

STc/ST and tsc

according to CHASE et al 1957

and

NORRBY 1972 for ships with

CPP.

>1, Vo>" 13 knots,

NI:

4<1 /Vo 4, 13 knots/

G 1, Vo < 13 knots.

% N (total

ob0

50

40

30

20

10

at>1

N89

0,2

0,6

0,8

SH/S

1,0

X4.1

N86 0,8

1,0

H/ST

Figure 58. Frequency distribution of SH/ST for ships with FPP and CPP,

DE and ST with

lie1and Voa`.. 14 knots.

50 40 30 20 10

0

3S

/V1/

3=

51

0

A±

0

AA

I A

A0

Ap

AI

AA

0

68

1012

1416

1820

Vkn

ots

A

A AA

A A

MY

&A

AA

Ae

0

1/3

Figu

re 3

9. S

L/ V

as a

fun

ctio

n of

Vo

for

X >

1, s

heer

S13

+ P

T. 0

:FP

P D

E, A

:PP

P ST

.

55.0

00

1111

1111

11.M

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iIIO

1111

1111

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1

Approach speed

17 knots

Full ahead

65.000

70.000

75.000

80.000

85.000

90.000

95.000

Displacement (Tons)

distance which can

be exceeded once in 100 crash sto s. 75.000 dwtHalf ahead

7 6Sl

ow a

head

5

2,0

F'5 0

55.000

60.000

65.000

siga

ll111

1111

1111

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1111

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0,10

1.1.

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=iii

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1111

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mm

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ftm

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amm

omm

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mE

mm

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mm

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mm

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==

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mm

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mill

sam

mm

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mm

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mm

o==

==

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iam

mam

mom

mw

mT

limnh

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limm

mm

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mom

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mm

ism

inum

milm

mam

mim

wiik

wii

mm

mm

mum

min

gim

mw

omm

mim

min

gsim

mm

imm

itnim

ffm

mm

umm

"mm

umin

ummusammummaimffiSWEENIMUMOMOSOMM

V N

E

15

14

13

90.000

95.000

70.000

75.000

80.000

85.000

Displacement (Tons)

Figure 41. Stopping distance which can

be exceeded once in 100 crash stops. 75.000 dwt

bulk/oil motorship, 17.600 BHT, 115 RPM,

app.

proach speed

17 knots

Full ahead

16 12

11

10Half ahead

8 5Slow ahead

1,5

EIM

ER

! E. M

EN

EM

.11

1116

1111

1111

EM

ME

E11

1111

1...m

.

mr"

"Pffi

ngIM

US

IME

NIII

IME

MO

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NE

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M11

1111

1111

1111

1111

1111

1111

1111

.

Witi

ffilff

ilWa

MlW

AH

MA

IVA

MIN

VO

NFA

VA

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AM

INFO

RE

M

NE

Eam

gmom

mE

NE

EN

IMIE

SS

MIT

EM

EN

IER

ME

MA

IMM

ICIE

Nw

aran

iam

mff

lum

mun

ikliM

MT

AM

MT

EM

BIN

EE

rmE

sior

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ME

MIL

TR

III

OsiirliFiiiEsTgaiseWinniriicur narientamosilrariewArgrar'Affiaurnirner.1"dmr

rRm

odui

alrm

wom

ulw

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n

TA

INW

ASE

I111

11

maw

=m

u na

ntem

te...

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---

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amom

mum

niem

ensa

mm

iNU

MM

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IIIIII

MIN

ST

INT

IMM

INE

WE

INK

IIMIM

EI

bral

illai

num

mom

aim

mum

muu

m-,

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rmaa

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mor

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illiESIIIIIMMINIESS11111111MINEM

EE

t:tt

t

roach speed

13

17 knots

12

-11

10 9 8 7 6 5

55.000

60.000

65.000

70.000

75.000

80.000

85.000

90.000

95.000

Displacement (Tons)

Figure 42. Lateral reach which can

be exceeded once in 100 crash stops. 75.000 dwt

bulk/oil motorship, 17.600 BHP, 115 RPM, FPP.

Half ahead

Slow ahead

m 1

,0

0 o 0,

5

cdP P

0

55.0

00111H

ER

ME

ME

NIE

MIE

ME

ME

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EM

INK

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HIP

MM

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ACKNOWLEDGMENT

It has been possible to conduct this investigation withthe goodwill and cooperation of the following personsand organizations to whom I wish to express my sincere

thanks:

E BOSTROM

CARLSSON

M CHIHAYA

A DELLHAG

C FALKEMO

R GETZ

B HANSEN

G HATTENDORFF

ILLIES

M JOURDAIN

S KANO

LACKENBY

LANGEVAD

T LEWIN

MEHLSEN

MENTZENDORFF

MEYER

S MOBERG

L MOSS

S MOTORA

T MOLLER CHRISTENSEN

NILSSON

H NORRBIN

A OOSTERVELD

W SCHMITT

SIMONSSON

J.STROMTEJSEN

I STAHL

Decca Navigator och Radar,Gothenburg

Ore sundsvarvet, Landskrona

Zosen, Tokyo News Service, Tokyo

Eriksbergs Mekaniska Verkstad,Gothenburg

Chalmers Tekniska HOgskola, Gothenburg

Skipsteknisk Forskningsinstitutt,Trondheim

Akergruppen, Oslo

Hamburgische Schiffbau undVersuchsanstalt, Hamburg

Technische Universitat, Hannover

Institut de Recherches de laConstruction Navale, Paris

MitsUbishi Heavy Industries, Tokyo

British Ship Research Association,Walls end

Decca Navigator, Copenhagen

Granges Rederi, Stockholm

Lindholmens Vary, Gothenburg

Technische Universitat, Hamburg

HowaldtswerkeDeutsche Werft, Kiel

Gotaverken, Gothenburg

Marcona Corporation, San Francisco

University of Tokyo, Tokyo

Odense Staalskibsvaerft, Odense

Kockums Mekaniska Verkstad, Malmo

Statens Skeppsprovningsanstalt,Gothenburg

Shell International, London

Rheinstahl Nordseewerke, Emden

Uddevallavarvet, Uddevalla

Naval Ship Research and DevelopmentCentre, Washington

Uddevallavarvet, Uddevalla

SOGAARD

T TANAKA

TANI

TANIGUCHI

I E TELYER

T WILSE

Nakskov Skibsvaerft, Nakskov

Mitsubishi Heavy. Industries,Yokohama

Tokyo University of MercantileMarine, Tokyo

Nagasaki Technical Institute,Mitsubishi Heavy Industries,Nagasaki

BP Tanker Co, London

Det Norske Veritas, Oslo

I would also acknowledge the help afforded me by friendsat BirdJohnson in Walpole, Baying in London, KjellbergKabushiki Kaisha in Tokyo and by colleagues on the staffof Karlstads Mekaniska Werkstad in Kristinehamn.

For the permission granted me to publish this reportthanks are due to Karlstads MekaniSka Werkstad.

CONTENTS

Page

Synopsis 1

Symbols 3

Introduction 8

Means and methods of improving the stoppingqualities of ships 10

Available data on board for estimatingships' stopping qualities 12

Methods for calculating stopping performance 14

Collection and processing of crash stoptest results 15

Course stability during the stoppingmanoeuvre OOOOOOOO 4 18

Sheering tendencies 23

The turning angle 26

Determining track reach and stopping time 27

Comparison of various calculation methodsfor estimating the stopping performance 44

The stopping equation for one per cent risk 49

Practical crash stop diagrams 51

Comparison of stopping performance with FPPand CPP ships 53

Summary 55

References 58

Tables

Figures

Acknowledgement