NEDERLANDS SCHEEPSSTUDIECENTRUM TNO
NETHERLANDS SHIP RESEARCH CENTRE TNOSHIPBUILDING DEPARTMENT LEEGHWATERSTRAAT 5, DELFT
NEDERLANDS INSTITUUT VOOR LASTECHNIEK
NETHERLANDS INSTITUTE OF WELDING
FRACTURE MECHANICS AND
FRACTURE CONTROL FOR SHIPS
(BREUKNIECHANICA EN BREUKBEHEERSING TOEGEPAST VOOR SCHEPEN)
by
PROF. IR. J. J. W. NIBBERING (State University of Ghent) (Delft University of Technology)
LEIT1D
In het kader van onderzoek naar meer verantwoorde
scheeps-constructies is reeds langer bekend dat de draagkracht van derge-lijke constructies voor een belangrijk deel door de vermoeiings-sterkte worth bepaald. In dit verband mag tevens worden gesteld dat de ontwikkeling naar meer gecompliceerde constructies en de toepassing van staal met hogere treksterkte de dominerende
invloed van de vermoeiingssterkte van de scheepsconstructie
sterker dan ooit naar voren doet komen.
Voor het ontwerpen van efficiente constructies is het daarom van belang het criterium voor een veilige draagkracht van een
dergelijke constructie als een verantwoord uitgangspunt voor
het ontwerp ter beschikking te kunnen hebben.
Een samenwerking tussen het Nederlands Instituut voor
Las-techniek" en het ,Nederlands Scheepsstudiecentrum TNO-leverde bijgaand rapport no. 178 S Fracture Mechanics and Fracture Control for Ships" op, dat in dit verband een
hand-leiding kan zijn.
Het rapport stelt duidelijk dat vermoeiingsscheuren als regel niet in het zgn. moedermateriaal van plaat en profiel ontstaan of zich voortzetten. Het verbindingsproces van deze basiselementen is meestal de hoofdoorzaak van de eerste aanleiding tot het op-treden van de vermoeiingsscheuren. De hoeveelheid laswerk aan de scheepsconstructie is van een dergelijke omvang dat kleine detecten die de scheurvorming kunnen inleiden, niet geheel zijn te voorkomen. Bovendien zijn de eigenschappen van het
basis-materiaal in de nabijheid van de las (heat affected zone) ten
gevolge van de met het lassen gepaardgaande warmtetoevoer, vaak beduidend slechter. Hierdoor zal een eenmaal geinitieerde
scheur zich in deze heat affected zone gemakkelijker kunnen
voortzetten en uiteindelijk aanleiding kunnen geven tot het op-treden van brosse breuk. Meer in het algemeen mag echter wor-den gesteld dat de plaats waar een dergelijke inleiding voor een scheur ontstaat van aanzienlijk groter belang is dan de afmetingen van deze scheur.
De toepassing van de in het laatste tiental jaren sterk
op-komende wetenschap breukmechanica" heeft veel bijgedragen
tot een beter begrip omtrent de praktische toepassing van de
vermoeiingssterkte van constructies voor geavanceerde schepen. NEDERLANDS SCHEEPSSTUDIECENTRUM TNO
In the scope of investigations into more justified ship structures it is already known that the capability of such structures will be
to a large extent governed by fatigue strength. The more so
because of a notable development to more complicated structures and the use of high tensile steels which cause a more dominating influence of fatigue strength on the capability of the ship structure. For the design of economic ship structures it is of utmost
im-portance that a reliable criterion for a safe capability of such
structures should be available as a starting point for the design. As a result of a co-operation between the Netherlands Institute
of Welding and the Netherlands Ship Research Centre TNO,
this report no. 178 S -Fracture Mechanics and Fracture Control
for Ships" may serve as a manual in this respect. The subject
report states clearly that in general fatigue cracks are not initiated and do not propagate in the base material of plate and profile. The welding process to join these basic elements is more often
the cause of the occurrence of fatigue cracks. The amount of
welding in ship structures is of such a magnitude that defects,
which can initiate cracks, can never be completely avoided.
Moreover, the good properties of the basic steel material in the heat affected zone has been more or less spoiled by the relative large heat input. Because of the latter a crack once initiated can propagate much more easily in the heat affected zone and ulti-mately may cause brittle fracture. More in general may be stated that the location of a defect which may cause a crack is of much more importance than the size itself.
The application of fracture mechanics, a science which developed rapidly during the last ten years, has given a much better understanding of fatigue considerations for advanced
structures of ships.
CONTENTS
2 Summary., , , page .General .. . ., . . ,...
- .5 1..1 Defects ., . .. .-.., :4 .2 li 4 '.-.. .. 1. ..;! , 5 1.2 Residual stresses ,..., .. . , 5 L3 Fatigue . . . .. .., .... ... - 6L4 High heat input welding . ,., , t 6 ,, , , . I
Acceptance testing and fracture mechanics for ships . 9
2.1 Crack-arresting ... . 9
2.2 Welding material . , A. ,A , 10
2.3. Heat affected zone, ..
. ...
102.4 The relative importance of metanurgicall factors and crack-length It
2.5 Critical C.O.D.'s and impact testing . II
2.6 C.O.D.-measurements . 14,
2:7 Final observations , 15,
Literature
APPENDIX: Introduction to fracture mechanics . . . 17
5
FRACTURE MECHANICS AND FRACTURE CONTROL FOR SHIPS*
by
Prof. ir. J. J. W. INIBBERING**
Summary
The amount of welding in ships is so large that defects can never be completely avoided. These defects will generally grow as a consequence of the cyclic loading a ship is subjected to. Mostly this does not endanger the situation; often it is improved because the quality of the material at the tip of an extended defect will be better than that at the original tip.
Thus the problem of fracture control in ships is more a matter of significance of defects in a metallurgical sense than in a
geo-metrical sense. In other words: location of defects and cracks is of greater importance than size. A small defect at a weld-crossing may be more dangerous than a large crack elsewhere.
Apart from these considerations it should be realized that ships are redundant stiffened plate structures. This has the following
consequences:
Brittle fractures, once initiated, may propagate to large lengths; 100% safety can only be obtained by using crack-arresting steels. Relatively large (of course arrested) brittle cracks and large fatigue cracks do hardly impair the strength of a ship as long as the involved reduction in sectional area is not too large. The steels used are so tough that crack length as such is hardly limited. It is obvious that limited room is left for usefully applying linear elastic fracture mechanics in commercial ships, except in the field of fatigue crack growth. General yielding fracture mechanics are more valuable, especially in terms of crack opening displacement (COD).
Notwithstanding this fact, in the appendix an introduction to linear elastic fracture mechanics has been given. Firstly because any structural designer should be familiar with some of the concepts, secondly because of the growing use of higher strength steels and steels for special purposes and thirdly because of the importance of fatigue-considerations for advanced ships.
Many points of this report are illustrated with results of experiments with Electrogas-welded plates of Nb-containing normalised St. 52. The harmful effect of heat-input both on the properties of the heat-affected zone, and on the residual stress field will par-ticularly become clear.
Indeed one of the purposes of this report is to give the necessary background for understanding a following report, in which low-cycle fatigue and brittle fracture experiments with 34 mm E.G.- and 46 mm ES-welded plates of St. 52 will be discussed. These investigations have been sponsored by the Netherlands Institute of Welding (NIL.).
1 General
Cracks in ships are either fatigue-cracks or brittle cracks.
Fatigue-cracks are most common but the danger involved is small.
Brittle cracks are very scarce nowadays, but when they occur, the ship obviously is in danger.
Both types of cracks generally start at defects due
to welding or flame cutting especially when these defects are situated at geometrical stress concentrations.
1.1 Defects
In ships with their enormous amounts of welds in
between 10 and 1000 km in length, weld defects can
never be completely avoided, despite intense non-destructive control.
The most important defects are under-cuts, lack of fusion, slag-inclusions, incomplete penetration and cracks in welds and heat-affected zones, (Fig. 5).
In principle one can achieve that not any defect will develop into a large brittle fracture. For this, it is
necessary that proper weld and parent materials are
chosen and welding methods are avoided, which excessively destroy the originally sound parent material in a relatively extended zone. Even the me,.e use of parent material of such high quality that any eventual brittle crack will be arrested, is mostly completely satisfactory. Then the quality of the welds and the H.A. Zones would be of secondary importance from a safety-point of view, and this could lead to substantial reductions
in cost of welding, production and quality control. Unfortunately when high heat input-welding methods like one pass submerged-arc, electrogas or electroslag-welding is applied, eventual cracks starting in the welds or heat-affected zones, are not always leaving the welded region under the influence of the residual welding stresses, as is the case with multipass-welding. (See 1.2).
1.2 Residual stresses
The stress field set up with high heat input welding has a much smaller gradient than the one created by low heat input welding, (Fig. I). Due to that the shear stresses in planes parallel to the weld are also smaller,
* Lecture delivered during the post-graduate course on Fracture Mechanics in May 1972 in Eindhoven. 9'. Professor Naval Architecture, State University of Ghent.
Reader Naval Architecture, Delft University of Technology. a.,
6 g.toad clod cc ----aload Fig. 2c.
Largest main stress deviates a with loading-stresses. :ventuat crack path goad Vain, stressh, Cres
Fig. 2a.. Fig. 2b.
Fig. 2d.
Largest main stress nearly in
line with loading-stresses (a
small).
and so are the corresponding normal stresses, of which the direction differs 45° from the weld line.
As fracture will always develop in a direction
per-pendicular to the main tensile stress, it will deviate more from the weld line, the higher the shear stresses be. In figures 2a and 2b the stress situation for a small element situated in the H.A.Z. is indicated; an estimate of the angle a between the. load stresses (nominal stresses) and the main tensile stresses is made in fig. 2c and 2d. It should be reminded that the shear stresses
Tres can only exist when the residual tensile stresses
(cres). also vary along the weld line. This is valid for hand-welding with many stops and' starts, but far less true for automatic and semi-automatic welding, especially with welds made in one pass..
Residual stresses have often beenblamed for causing brittle fracture. This was right for ships constructed during and shortly after World War II, when low-stress brittle fractares were very common and the steel was often not good enough to arrest them. But nowadays residual stresses are rather beneficial. The ships' steel. is so much better than formerly, that eventual cracks started in a weld or H.A.Z. are mostly arrested in the parent plate before attaining a critical length. (This is not the critical length under static conditions, which
!It
is in the order of meters, but the critical length be-.
longing to the high-speed state of loading occurring when a brittle crack propagates).
1.3 Fatigue
The residual stresses have also a great influence on the propagation of fatigue-cracks: partly beneficial, pary detrimental..
The beneficial effect is again that Cracks are forced to leave the welded region. The detrimental effect may be that cracks initiate and propagate easier when
residual tensile stresses are perpendicular to the crack front. This is generally the case with cracks developing perpendicular to a butt weld. In other cases, for
in-stance when cracks develop more or less parallel to a
ir
butt weld, the propagation may be slowed 'down. f course the heterogeneity of the H.A.Z. also play a
role.
All these influences make the use of a simple.crack propagation formula like
= c(z1 Kr (see appendix) dN
very unreliable. The value of in can vary between 1.5 and 6 [1], For shipbuilding, problems like the random character of the loading, combinations of in-plane and normal-to-the-plane loading and corrosion add to the' complexity of the problem.
On the other hand it is fortunate that a high accuracy in estimating fatigue crack growth is not important
for ships. One can hardly speak of a critical crack
length, because the notch toughness of shipbuilding steel is good enough to be able to bear the presence of really large cracks. Leakage therefore is. a more im-portant restricting factor. But as most cracks develop in internal stiffening members, leakage often can be neglected.
It has been emphasized that due to the inevitable
presence of defects, cracks soon develop, under the action of the cyclic loading and consequently the re-sistance to propagation of cracks is mainly determining the fatigue life of ships' structures. Nowadays this is even more true, because of the very large dimensions tankers, bulk carriers and containerships have obtain-ed. A crack of 10 cm .depth in a 50 .cm high stiffener If a small ship reduces the cross section as much as a 40 cm crack in a 2 m high stiffener in a large ship. Yet the time for development of the latter is about
two times as long as for the former.
Fatigue-cracks are not only relatively harmless Such, they can even be beneficient.. For, whenever
brittle cracks develop in ships, they always start at points where the material is in an extremely
bad-condi-= Cre ii,' r.' '.;. i -...r
'
Eventual crack pathLOW HEAT-.INPUT HIGH HEAT-INPUT
Fig, 1.. ITres. Tres. 0-brad as Cres. Tres tires Cioad
tion, and this of course is often the case at the tip of weld- or H.A.Z.-defects (see section 2). A little exten-sion of these defects will shift the tip from material in had condition to that of normal state, and initiation of a brittle crack becomes far more difficult,
IA High heat input welding
There are however some points which need special
attention. Firstly again the case of high heat input
welds. It is very well possible, and demonstrated in [1], that fatigue-cracks propagate completely in the weld or H.A.Z. (see Fig. 4). In the case considered a
I ig,. 3a. Extreme grain coarsening in H.A.Z. of E.G.-welded fine grain steel, (Nb-containing normalised).
28
24
20
16
25 35
3 mm wide (or narrow!) coarse grained zone existed of very poor notch toughness along an E.G.-weld in a
34 mm plate. Notwithstanding the narrowness of
this zone, high stress-low cycle fatigue cracks indeed propagated in that zone over quite a distance (up to 120 mm). The large danger involved can be appreciated when it is known that the fatigue-crack "jumped" forth in a brittle way over a small distance several times
(5-10 mm) before it finally developed into a complete brittle fracture which also ran all along the weld line! The test-temperature was 20°C, the nominal fracture stress 24 kg/min'. The steel was extremely good (E-quality), as
is obvious from the Charpy V notch
-20 -10
PEAKTEMPERATURE
0
-80 -70 -60 -SO -40 -30
TEST TEMPERATURE ( °C )
Fig. 3b. Charpy V values of the H.A.Z., St. 52+ Nb (thickness: 34 mm) and the E.G.-weld151.
0
PI_ ATE MATERIAL
rIrw
0A14
The distances from V-notch to fusion Line and the peak temperature at that place during welding are indicated
at the curves. sri
Plate material (unaffected) --- "4,;;Ciiii ---+10 +20 +25 51500 1F0SiONLINE 12 DISTANCE FUSIONLINE 35mm 6mm(1020%1
grra....---8 FRACTURE SURFACE-FRACTURE SURFACE FRACTURE SURFACE,t FRACTURE SURFACE sSIOMN- LINE NOTCH BRITTLE STEPS NOTCH E.G. -WELD .G -WEL ifDETAIL NOTC .4k
Aiolvigii"- - abk%*to ..4,4titsar
NOTCH 2b
MN>
E.G.-WELD 'FUSION- LINE ° 4 - _."_ -NOTCH SA. ELD FRACTURE SURFACE CRACKED AND NOTCH ARRESTED NOTCH 2aFig.. 4. Crack paths in E.G.-welded axially loaded plates. (For specimen see figure 8).
FRACTURE PLANE SPECIMEN, 1 NOTCH. 1F 1 FIG. N2. FRACTURE PLA E SPECIMEN, 2 INOTCH. 2F1 FIG. NT FRACTURE PLANE SPECIMEN. 12 'NOTCH. 2*2:(12),
E_G)_- WELD) am s
'SIGN-LINEMINIM
FRACTURE PLANESPECIMEN. 12' NOTCH. 2eof FIG. NB. 11, FRACTURE PLANE SPECIMEN. 2' NOTCH. fa AND 2.5 FIG, NS. FUSION-LINE E.G-WELD. 1DETAIL S.A 1
energy which was 16 kgm/crn2
at 60°C! The
reduc-tion in quality in the H.A.Z. due to the high heat
input of the E.G.-welding is shown in figures 3a 'and .3h. The crack paths are shown in figure 4.
In terms of "safe" temperature the steel has been
spoiled nearly 100°C. Similar dangerous situations
can also occur with more normal welding methods - especially when the number of passes is not very high or when one side butt welding has been applied.
Figure 5 shows some examples,
ti Incomplete
epenetration
Section weakened by
Fig., 5. Main weld defects.
It will be clear, that 'cracks developing at the points of incomplete penetration will not easily leave' that
vertical plane, as long as the cross-section remains reduced by that lack of penetration. A similar situation may occur at undercuts and long lacks of fusion, (Fig. 5).
It is evident that transverse butt weld's in ship decks, sides and bottom are more critical in this respect than longitudinal butt welds. But generally speaking only reallylongdefects are dangerous. Most of the standards and specifications as to defects are too pessimistic. On the other hand it is well recognized that these specifica-tions have mainly the purpose of guaranteeing a satis-factory level of workmanship, thus having a function in quality control which of course can never be dis-pensed with.
2 Acceptance testing and fracture mechanieg for ships
The currently used ships' steels and welding materials
have such good toughness that the application of
linear elastic fracture mechanics and the use of lc. or
GI, as minimum required toughness values for static
Slag
'Line' stress
concentration acts as crack
track.
loading have little' sense. 'This is partly due to the
severe requirements set for these materials, especially because not static but impacttests. are prescribed. The second cause is that the temperatures ships are sub-jected to are not too low, connected to the fact that the
lowest sea temperature is about -4°C. Finally for ship steels metallurgical influences, due to welding and flamecutting overshadow largely the influence of geo-metrical factors like crack-length (see 2.4).
Notwithstanding this the meaning of K and some principles of classic fracture mechanics are illustrated
in the appendix in order to make the reader a bit
familiar with these Furthermore their limitations forshipbuilding steels are explained.
2.3 Crack-arresting
When it is required that ships' steel in deck. and bottom should be able to arrest any brittle crack, fracture
mechanics can be of some use. But the Most realistie way for estimating the involved dynamic toughness of the material (Kroc.) is by carrying out a crack-arrest
test of the Robertson type, and evaluate it for the
maximum length or crack to be arrested. 'That length May be appreciable because a crack started in the side plating, may have developed many meters before it finally meets the sheer strake in which it should be arrested. However, when looking to the usual results
obtained from isotherm Robertson tests it always
OK ARREST THROUGH CRACK
it CRITICAL X N.C.R.'E.
.0 STEELMAKERS
CONSTANT TEMP TESTS{N'C'R'E'
ST EE L MAKElqS
6 FT. WIDE PLATE TESTS.
A ft
A
IKt
, I X .X "0il
6,1 ± ii I -100 -80 -60 -20' Temperature (°C),Fig.. & Results of crack-arrest tests (from admirality report RD.. 20 16 12 tri 2 'E' 4 s/t 40 0
shows a tremendous increase in dynamic toughness over a very short temperature region; for instance 5-fold KID, over 10''C, (Fig. 6).
As K, is about proportional to the square root of crack length, a 5-fold increase in K1, would for a certain temperature mean a 25-fold increase in critical crack length. In other words: the sensitivity of the
material to temperature is much larger than to crack length.
In quantitative terms: when a material has an iso-thermal Robertson crack arrest temperature of 10°C which means that cracks up to about 30 cm can be arrested then it will be able to arrest cracks of 5 m-10 m length at 0C. In such a situation it is of course wise to require a Robertson crack arrest temperature
of say 15°C and assuming it may be 5°C too safe. Most steels of D and E quality are certainly more "safe".
It is unfortunate that in the specifications of steels reliable crack arrest tests are not required. The desired properties are now more or less obtained by requiring certain Charpy-energies at certain temperatures. The correlation between these transition temperatures and
the Robertson crack arrest temperature is not only
poor, (the latter is on the average some 40°C higher than the 3.5 kgm/cm2 Charpy temperature), but the scatter is also large. For individual steels the difference between 3.5 kgm/cm2 Charpy and -Robertson crack
arrest temperature can be 0°C to 80C. (See Fig. 7,
obtained from [2] by Drs. H. C. van Elst). This does not mean that the Charpy-test has no use anymore in this respect. Verbraak and Van Elst have already long
ago propagandized to use it for quality control of steels with specific composition, made according to a
fixed procedure. For such a well defined case, the
relation between the Robertson crack-arrest temper-ature and for instance the Charpy-energy at that tem-perature shows little scatter and can be used for quality
60
0
20 o °/.,
100 210 6 20 40 60 80 3,5kgm/cm2 (.20 ft Lbs.) CharpyV temperature (°C)Fig. 7. Relation between 20 ft. lbs. charpy-transition tempera-ture and crack-arrest temperatempera-ture (Van Elst [21).
control. In the same way the Charpy test can be used
for quality control of welding material.
Notwith-standing this a more realistic and sophisticated yet simple test like Pellini's Drop Weight Test is preferred to the Charpy test (see 2.7).
2.2 Welding material
In the I.I.W. the W.G. 2912 has the task to develop
reliable quality control and acceptance tests plus criteria for welds. During the chairmanship of Van den Blink, the group arrived at the opinion that for weld material crack arresting is of no importance, because as mentioned in section 1, cracks always tend to leave the welded region.
That opinion was based as well on overwhelming practical experience as on results of experiments from Kihara and Ikeda [3].
The conclusion was that for welds, only the resist-ance to crack initiation was of importresist-ance.
If so, it would be realistic to stipulate less severe requirements for welds than for plate material, because the latter have to be resistant to crack propagation
which is a severe dynamic phenomenon. Of course initiation of cracks can also occur as a consequence of dynamic loading, but seldom as severe as occurs
during brittle crack propagation, of which the speed is some 2 km/sec. But it will be understood from the former discussions on the reduced influence of residual stresses in case of automatic and high heat input
welding, (section 1.4), that the original opinion of the
W.G. 2912 has to be reconsidered. Propagation in welds has become real.
2.3 Heat affected zone (H.A.Z.)
Another point is, that not always the H.A.Z.-material properties are controlled; yet they may be worse than those of the weld metal. The inspection boards meet
this by having a built-in safety margin in their (Charpy) requirements for plate material which allows for a certain deterioration of the plate material
in the
H.A.Z.One might consider it in this way: when the plate material is resistant against brittle crack propagation,
the H.A.Z. will at least be resistant against the much milder crack initiation. This method may be right for normal, multi-pass welded steels; it is wrong for nor-malised or quenched and tempered fine grain steels, welded with high heat input.
This has already been demonstrated in figure 3. A logical approach would be by taking many Charpy-specimens from all over the H.A.Z. and stipulating a lowest acceptable value. This indeed occurs more and more, but unfortunately not always realistically. For instance 6.2 kgm/cm2 at 30°C as required by some
10
0
°
inspection boards for ships, may be much too severe. In the Delft Ship Structures Laboratory full scale experiments with 34 mm thick E.G.-welded plates of
Nh-normalised steel have proved that a minimum value of 1 kgm/cm2 Charpy-energy at the lowest
service temperature would be sufficient for the case considered (see [5] and section 2.4 in connection with
Fig. 3b).
2.4 The relative importanceofmetallurgical factors
and crack-length (see also appendix figure 18)
In section 2. I on crack-arresting, it has been discussed that the influence of temperature on the toughness of a material subjected to high speed loading is so large, that crack length is only of second rate importance. When discussing crack-initiation, the situation is
similar in very brittle conditions. These conditions
may exist in Heat Affected Zones. As an example again is chosen the H.A.Z. of an E.G.-welded
Nb-normalised thick test plate.
In figure 8 the specimen is shown, and in figure 9
some results are given. From the numbers near the points it can be seen that crack length had not any influence on the results. But the distance from the
crack-tip to the fusion line was of utmost importance: a difference in position of 1 mm had a greater influence than a difference in crack length from 26 mm to 100 mm. In figure 9 both influences do not appear as clear as possible. A presentation of C.O.D. (see 2.6) as a function of temperature, as in figure 10, is much more discriminating than fracture load versus temperature. Following a vertical line in the region of 25°C crack lengths increase from 26 to 36, 53, 76, 104 and 79 mm's. This is quite opposite from what fracture mechanics would predict. But another look shows that the lowest points generally belong to cracks situated on, or close to the fusion line (s/s) and the higher points to cracks further away. Together with figure 3 this demonstrates the large differences in material quality in a very
narrow region. The figure also shows the enormous
notch"B"4
lT5
600 600
Fig. 8. Specimen loaded in 1000 ts testing machine [5].
600
influence of temperature on the C.O.D. to fracture: 10°C rise in temperature raises the C.O.D. from 0.015
1/3 .
to some 0.5 mm, (point figure 10).
100
2.5 Critical C.O.D.'s and impact testing
In such cases of course it has little sense to discuss thoroughly which value of C.O.D., should be required as an absolute minimum. Whether it is 0.1 rum or
0.2 mm does not make a real difference in structural quality. On the other hand, for automatic welds much smaller increases of C.O.D. with temperature have been observed (Fig. 11 [6]). Therefore it has been
proposed to require always at least 0.3 mm C.O.D., in static testing.
This will only be in exceptional cases too much on the safe side, but for most cases it will be as good as a slightly different value. It has however the advantage of being easily measurable.
It should be underlined that static C.O.D.-testing is not always, or even seldom, sufficiently realistic.
Many structures may be subjected to regular or accidental shock loading.
In view of this it is in a way fortunate that most
inspection boards define their specifications in terms of energy obtained from impact tests.
What difference this makes for weld metal is evident from table I, column 3 and 5.
Table 1. [6]
i150
IIPI i1 1notchA" notch "F1" y V vecaotteccerz..saverstmemorwovimarvir vityl V .t, S A-weld V'
i notch "F2 " 1®
i
5.
E 600 150 notch"EA. La) ' 'notch "G" notch "0"1100
sooNiblink test C.O.D.A. test
Charpy dynamic static
3.5 kgm/cmz 0.06 mm 0.3 mm +10°C +20°C +22°C 4°C
46°C
+13°C +18°C19°C
+13°C10°C
8°C 3°C62°C
18°C
13°C
23°C
+73°C +27°C24°C
+ 5C
63°C
WELDS® (...5D (6)AND 7 ARE ELEC TROGAS WELDS IN 34 rnm PLATES AND EL ECTRO SLAG IN
WELD ® SUBMERGED-ARC WELD. 46 mm PLATES.
THE NOTCHES C, "F1" AND "F2" WERE MADE AFTER WELDING
® AND ® AND BEFORE WELDING ® (61.W - NOTCH ). THE OTHER NOTCHES WERE MADE AFTER WELDING.
300 thick-type of ness weld in mm ES 46 Sub.Arc. 46 EG 34 Sub.Arc. 34 EG (A) 22 Sub.Arc. (A) 22 EG (B) 22 Sub.Arc. (B) 22
12 42 41 40 39 38 37 36 3 26 0.5(N+Wl. 26 .1/31_t 26 1/3 26 6/4 75
S A -weld/ 3(Tensile test after fatigue-Loading)
101
1/3 100
itrjr13/3
31 (Tensile test after fatigue- load ng )
PLATE 26 3/3 27 HIGH STRESS FRACTURE 8/6 68 X ST 52.52 + Nb. E.G.-welded. (Thickness: 34mm) SPECIMEN 1 (n.34000 ) 0 2 (nO) NO + (n1500 FRACT-URE
"
"
12 (n.660) 1:1 12' )n.5800) 91 COMPLETE FRACTURE NUMBERS AT POINTS. DISTANCE TO FUSION-LINE AND TOTAL CRACK-LENGTH ( N+W ): NOTCHED +WELDED20 i i
i iI
I-27 -26 -25 -24 -23 -22 -21
-20 -19
-18 -117 -16 -15 -14 -13 -12Test temperature (°C)
Fig. 9. Net fracture-stress as function of temperature [5].
-11 34 33 32 8/61 65 26 31 533 ('s1 30 LOW STRESS .1141 FRACTURE E31 29
28.0
t/3
26 c .36 027-
41/326 U CD 3/3A 0 I 26-26 I + _c U 2.5 25 i-( N+6.. NOT FATIGUE-LOADED 26(SEE FIG 1 (SEE FIG 21)
i PLATE 24 2 10/0 0.5 ( N.v., l:6 26 0.5(N+It 26 11/3 I I50 26 23 22 21
+10
26 1/3 79 2.' I > 27 I - - 1 8/680 70 60 50 40 301 20' 18 16,1 14 12 10 9 t_ gm. VI 791 S.A,-weid/
(TENSILE TEST AFTER FATIGUE -LOADING )
104 SEE
I
6/41 76 1/3 * 1/3 36 1 2.5(N+w) 26 (SEE FIG. 11'3 Vs° 26 .4W 26t3
5350
(Not fiatigue-loade4 t 01601/
SA.-weld/
64/
75 (SEE FIG. 2$, ,,../. 2X.5 1 0.5 (N+W) /26/
1/3 42 3/3 44. 3 26 18 6/4 0/0 54 PLATE 26 Si ND FRACTURE :011(N+W)OT°... PLATE 26 .44.(n10000) '2 '13/6 5.433/:.( TENSILE TEST AFTER FATIGUE-LOADING)
/661
ST.52 +Nb. -welded Thickness: 34 mm. SPECIMEN 1 )n:34000N o 2 (n.,0'))2' (n41500) X, 12 (n.660) a 12')n5800) COMPLETE FRACTURE lg. PARTIAL FRACTURE 65 0.5 ISPECIMEN. 1 ( 50 NUMBERS AT POINTS: DISTANCE TO FUSION-LINE AND TOTAL CRACK-LENGTH.'12/6IltliC lItt4 1 3 26 41N +W ): NOTCHED *WELDED 1, 01 I. gr il 6 1 1 1 1 - I. - I - I 111 II 1 II -28 -27 -2,6 -25 -24
-23 -22 -21
-20 A9. -18 -17 -16 -15 -14 -13 -12 -11 -10Test temperature (IC)
Fig. 10, Deformations at the notch's tips at the itioinent the fracture occurred given as function of temperature 1[5].
7 6 5 4 3 ( 100
9
E 0 1 2' 1 1±1 5 8 2 0.51
0.
15
0,3
-U2,,
Test specimens according to can& committee plate thickness
-
46 mm (E.S.' 34 mm (E.G.) 22 mm(A)(E.G3 --- 22 mm (3)(EG ) 0 (8) .7..22mm 46mm 4.. 0Fig.. II,. C.O.D. at fracture versus temperature (electro-slag and electro-gas welds),
ry
Although impact testing especially for shipbuilding Fig. II The C.O.D. concept. material may be considered to be realistic or at least
(1) Becomes: on the safe side", the ,Charpy-test is not very fit for
this purpose., hx.=4a .\/(a+r,)2. x2
The main shortcomings of the Charpy-test are,:
a. reduced plate thickness; b. small, rather blunt, notch;
c.. recorded energy is sum ofinitiation andpropagation
energy;
d., instead of energy, thedeformationat the notch root This results, irt:
(C.O.D.) should be measured. E
These objections have been avoided in the so-called = 40 Niblink-test [8], which is actually studied in the 1.I.W.,
W.G. 2912. For information, some
given in table I..
2.6 CI 0 .D.-iireastirements
The attractiveness of the C.O.D.-concept is that its
physical meaning is quite clear: it is the deformation
(extension) of the material at the tip of a notch or
crack. Wells [7] introduced this concept.
ness mainly to their large plastic deformability, so that a ductility-requirement is a logical consequence.
As discussed before. critical COD's for these steels should be in the order of magnitude of 0.2 to 0.4 mm. Even for cracks as large as 100 mm in length, these values can only occur at nominal stresses equal to or larger than yield point.
This can be illustrated as follows (see Fig. 12). For a crack with a plastic zone 2ry,the virtual crack length is about 2(a
For a crack without a plastic zone applies
a
A crack with -a plastic zone has a virtual length
= a+ ry
PLzstic zone
For x = 'is.
40
(5-x
- (a+
ry)= C.O.D. = ry2+2ary,
E (relation 6, p 4-+ry).
Substitution (3) in (4),
COD =
For ilower strength steels of moderate thicknesses it is much more realistic to specify required fracture
toughnesses in terms of critical C.O.D.'s than in terms
of K1 or G. These steels owe their structural useful- (Relation C.0.D.4 6,)
111
(1)
0)
results are also so the plastic zone can be calculated when6, has been measured. Substituting x= a in (2), ) (5) 10 -70 -60 -50 -40 -30 -20 -10 0, +10 +20' Test temperature MI 1,3 1,2 -1,1 1,0 -0,9 0,6 r E a' ry 0 = = a (3) =
From
a Va2 x2
E
the C.O.D. at crack tip can be calculated by sub-stituting a+ryfor aand afor x
C.O.D. = V ry2 -1-2ary;
E
when C.O.D. = 0.4 and o- = 30 kg/mm2 ry=-35 mm. When linear elastic fracture mechanics would be valid for this situation, the nominal stress should be equal to the one calculated from
(see appendix)
au2
ry=
2u;
For a steel with a yield point of 35 kg/mm 2 this leads to
This is larger than yield point indeed, and linear elastic fracture mechanics do certainly not apply.
For a C.O.D. of 0.3 mm, ry= 22mm and uca,c = 33 kg/mm2. This is also too close toay for valid calcula-tion results.
One disadvantage of C.O.D.,. as a fracture criterion is, that it is probably not independent of crack length.
For from
C.O.D. = ry2 +2a- ry
it can be seen that when ci o-, the larger the crack, the larger the C.O.D. at a certain size of plastic zone. Now, as is known [7], the size of the plastic zone (in relation to plate thickness) is a measure for the stress
state in that plastic zone, (ry t fully plane stress (see Fig. 15)). Consequently the larger a crack the
larger the C.O.D. needed for plane stress.
Nevertheless, in view of the many unknown factors involved, and for cracks or defects between 10 and 100 mm length, one single C.O.D.-value will suffice. It has the advantage that for very thick plates, Wells' requirement of ry I does not lead to unrealistically
high toughness values. A C.O.D. of say 0.3 mm indeed represents a quite satisfactory deformability. The fact that it corresponds to a situation of plane strain
in very thick plates is not objectionable at all.
2.7 Final observations
Figure 10 is a good illustration of the usefulness of the C.O.D.-concept. There is a clear separation between the "bad" and "good" results, which contrasts sharply
with figure 9. The figure however demonstrates also how important welding parameters may be, and that these may completely overshadow the influence of other parameters like crack length. This has already
been discussed before as far as the position of the
cracks relative to the fusion line concerns. But in this final section the attention is drawn to such an apparent-ly secondary factor like sequence of welding. The point quite to the right:
0 (N+ W)
28; n 10,000
represents a partial fracture, which started at a notch, which was made in the H.A.Z. of a transverse
E.G.-weld prior to the subsequent longitudinal submerged
arc welding. The distance between the notch and the S.A.-weld was so small that the material at the notch tip during the S.A.-welding was strained plastically to and fro at a temperature between 300 and 500°C. This caused hot straining embrittlement, which once more reduced the quality of the H.A.Z. at the tip concerned appreciably. (Greene-Wells' embrittlement).
A similar effect may occur at weld defects close to weld-crossing (Fig. 13). The transition temperature of the transverse (first) weld can be increased some 60°C, when the defect
is at a critical distance from the
(second) longitudinal weld.
Fig. 13. Defect at weld crossing.
The unfortunate dilemma for the structural engineer is, that good engineering practice is a welding sequence as indicated in figure 13. For then, there is the smallest chance that weld defects occur. But whenever they occur, the situation may be far more dangerous than when defects are present in alast made transverse weld. For cases like the former, two solutions are available: Non destructive testing of all weld crossings, (also where stiffeners cross butt-welds).
Require that the weld metal and H.A.Z. have a
u = 2ry-o-y2 a 2 35 352 50 = 41 kg/mm2 y 2..
16
PeRini D. W.-transition temperature 'rower than service temperature. There is a great similarity between the initiation of -a crack at an embrittled
defect tip and in a notched brittle weld as in
Pellini's Drop Weight Test.
Literature
1 NIBBERING, J.. J. W. and A. W. LALLEMAN, Low-cycle fatigue
problems in shipbuilding; crack propagation in coarse-.
grained zones of thick plates. Proc. Conference on fatigue
of welded structures, Brighton, July 1970,, paper 16.
2. ELST, H. C. VAN, Over het onderlinge verband tussen brosse breulcproeven met kleine en grote proefstukken. Lastechniek no. 8, 1967.
3, KIFIARA, H., Recent studies in Japan on brittle fracture of welded steel. 11W-dm. X-291-63.
4. A comparison of transition temperatures determined by
small and large scale tests on five steels. Adm. advisory com-mittee on structural steel. Report P2 1960.
NJIBBERING, J. J. W. and A. W. LALLEMAN, Low-cycle fatigue tests at low temperature with E.G.-welded 34 mm plates of St.. 52 NI). 11W-doe.. X-593'-70.
NIBBERING, J. J. W., Comparison between static COD-tests and Niblink drop weight tests 11W-doe. W.G. 2912-168-72.
7. WELLS, A. A., Application of fracture mechanics at and
beyond general yielding. British Welding Journal, Nov. 1963.. & BLINK, W. P. VAN DEN and NIBBERING, J. J. W., Proposal for
the testing of weld metal from the viewpoint of brittle fracture initiation.. Report NSS-TNO no. 121 S Oct. 1968.
APPENDIX
Introduction to fracture mechanics
Most introductions on fracture mechanics start with energy-approaches.
The strain energy released at a unit extension of an existing crack (e.g. extended by saw-cutting) is called G. At the moment the released energy becomes larger
than the energy the material close to the crack tip
can absorb (by deforming) an unstable fracture starts.
In other words: the strain energy release rate has
become critical (Gc); the resistance of the material to crack extension was no longer sufficient. This resist-ance often called the fracture toughness is the same as
2 a
2b
a(r, 0) = ci xf, (a, b, etc.) xf2(r, 0)
load notch position of geometry a(r, 0) Stress intensity =f3(r)xf4(0) = parameter K f= x 4(0) 2nr for 0 = 0 is f4(9) =-- I .f2(r, 0) = = 2nr I g 1 N/2.irr
For a through-thickness notch in an infinite wide plate is:
(TN/ na
Ia
K = xf i(a) =
a =\ 2nr v 2r
Fig. 14. The stress intensity parameter K.
It can be shown that
G =-maa2,
in which: o- = nominal stress a = half crack length E = Young's modulus.
The foregoing has only been mentioned for the sake of completeness. For there is a more simple way to attack fracture problems by only looking to what happens at the tip of a notch or crack.
But we should forget thinking about stress concen-trations. They depend too much on the sharpness of the notch and on the amount of plasticity occurring at the notch's tip. It is much simpler to say: what happens exactly at the tip is not feasible, but it is determined by the state of stress and strain in the immediate vicinity of the tip. When we are able to characterize that state we have obtained a very efficient tool for solving crack problems.
The stresses a(e) near the crack-tip depend on:
I. The nominal stresses (a) cr, =
c,-2. The geometry and position of the crack.
For a centre-line through crack perpendicular to uni-axial load stress (Fig. 14) crack length 2a is the only geometric parameter that counts.
It can be shown that o- r, = c ,N/ a.
When 1 and 2 are combined, the stress field around the crack tip is completely defined by a-\/ a. Historically na has become usual and has been designated by K. (For other cracks, e.g.: part-thickness ones, instead of 1 na other expressions are valid).
For pure elastic problems, where linear elastic fracture mechanics apply K suffices and ar,,, is not of any interest. But in cases of small scale yielding we need expressions for the size of the plastic zone, and for this we need to know the actual stress distribution. It can be shown that
ar,0 = __f (0)
,2rcr
For
= 0,f(0) = I = ,K
N/27.rr
The use of K will be explained a little more. When a tensile test is carried out with a plate as in figure 14, K will rise in proportion to the load (K = na), as
long as no extensive yielding occurs. (This will be
discussed further on). When a fracture develops the nominal fracture stress might be called the fracture
-Gc...
a =
strength of the plate. But it is not a measure for the
fracture toughness of the plate-material. For with
another initial notch, a different fracture strength would have been found. Consequently fracture stress
is not a material property as in the case of an
1.117-notched bar. A good material property is lc, the
K-value at the moment of fracturing. For, one will
observe from the test results that a1N'2Ia1 of the first plate is equal to o-2\ Itra, of the second one (al and a2 are the nominal stresses at fracture). In other words: fracture develops when K reaches a critical value lc which is called the fracture toughness.
Nowadays K and lc have largely replaced G and
Gc mentioned in the beginning of the appendix, the
relation between both is K2 = E.G. (plane stress) or K2= E.G./1 v2 (plane strain), see figure 15.
SMALL PLASTIC ZONE; MATERIAL AT CRACKTIP CANNOT CONTRACT FREELY, BECAUSE IT IS SURROUNDED BY NON -PLASTIC MATERIAL.
111111111111WASW
PLANE STRAIN
LARGE PLASTIC ZONE ( )
MATERIAL AT TIP CAN EASILY CONTRACT FROM t to
111111111111
111111[11111112
PLANE STRESS
Fig. 15. Plane strain and plane stress.
Instead of Kc, the symbol K10 is often used. The
subscript refers to fractures perpendicular to the notch plane like cleavage fractures in the case of steel. For all materials the fracture toughness is a function of thickness. The larger the thickness the smaller Kc. At a certain thickness a minimum lc is found which is called
K:
the plane strain fracture toughness. In fact only this value is a real material property. (For lower strength steels only at a particular temperature). Apart from the large capacity for plastic deforming,and the temperature dependancy of mild steel, the
strong influence on it of fabrication procedures like cold forming, welding, flame cutting is of utmost
significance. It makes that K-values for unimpaired parent metal have little practical value (see figure 18). One might object that they could be useful for
estimating the crack-arresting ability of a steel. But
then another fact should be realized viz, the large
dependency of Kfc on loading speed. For
crack-arresting only K,,,.-values are of interest, (D =
dy-namic), but in most cases the so called crack arrest
transition temperature has a greater practical value, (see Fie!. 6).
Correction for plasticity
For most industrial materials the initiation of a crack is preceded by more or less plastic deformation at the crack tip.
As long as the plastic zone is small (in relation to
plate thickness) the stress state at the crack tip is
strongly triaxial (plane strain). Then the average yield point in the plastic zone is about N./3 x The size of the plastic zone can be calculated by realizing that at the edge of that zone (at r =r, in Fig. 16) the stress is equal to the local yield point. But that is not sufficient, it should also be taken into account that in the elastic (full line) and the elasto-plastic (dotted line) situation
the stresses should be equivalent. It can be shown
that this results in a size of plastic zone S to be equal
to 2r.
Fig. 16. Plastic zone in plane strain condition (aYtocal = '/30Tnorrnal; gy= yield point).
As said before, rycan be calculated.
In
ar =
7rr
one should substitute ay,/.3 for and ryfor r.
K2
r =
61to-2
For plane stress
Cr -= Cry y =
2tra-2
In the plastic condition the stresses are cut off to the
level of the local yield point, but the strains keep obeying more or less the formula:
E = (E- Er now replaces ar).
127cr
2a
v-3-.
(plane strain).
Elastic stress distribution Gry Plastic stress dir distribution 18 r cryie,d. K2
in"pLane strain" condition:Vi.Gy
in"plane stress" condition: Gy
It
v.Non-real. part of the
stress -function in the
\ elastic region.
Crack Length for
calculation of K
Fig. 17. Virtual crack-length in presence of plastic zone.
Fracture becomes rather a consequence of exhausting
the capacity for deformation, than of surpassing a certain critical stress-value.
K should be corrected for the presence of a plastic
zone. In figure 17 (from [7]) is illustrated that K is larger than in the fully elastic condition, because a
notch with a plastic zone has a virtual crack length larger than the notch's length K = n(a+ry).
The K-concept is of great value for fatigue-problems, (see 1.3), and fracture problems of very high strength materials. With steels of moderate strength fracture practically always occurs at nominal stresses close to or exceeding yield point. It will be clear that then the
K-concept has not any significance. It should be replaced by the C.O.D.-concept. C.O.D. = crack opening displacement at the notch's tip: it is discussed in section 2.6.
Crack Length Crack tip position
Fig 18. Relative importance of crack-length and crack position in welded region.
-PUBLICATIONS OF THE NETHERLANDS SHIP RESEARCH CENTRE TNO
LIST OF EARLIER PUBLICATIONS AVAILABLE ON REQUEST
PRICE PER COPY DFL. 1()- (POSTAGE NOT INCLUDED)
M = engineering, department S shipbuilding, department C = corrosion and antifouling department
Reports
90,S Computation of pitch and heave motions for arbitrary ship forms. W. E. Smith, 1967.
91 M Corrosion in exhaust driven turbochargers on marine diesel
engines using heavy fuels. R. W. Stuart Mitchell, A. J. M. S. van Montfoort and V. A. Ogale, 1967.
92 M Residual fuel treatment on board ship. Part II. Comparative
cylinder wear measurements on a laboratory diesel engine using filtered or centrifuged residual fuel: A. de Mooy, M. Verwoest and G. G. van der Meulen, 1967.
93 C Cost relations of the treatments of ship hulls and the fuel
con-sumption of ships. H. J. Lageveen-van Kuijk, 1967.
94 C Optimum conditions for blast cleaning of steel plate. J.
Rem-melts, 1967.
95.M Residual fuel treatment on board ship. Part I. The effect of cen-trifuging, filtering and homogenizing on the unsolubles in residual fuel. M. Verwoest and F. J. Colon, 1967.
96 S Analysis of the modified strip theory for the calculation of ship motions and wave bending moments. J. Gerritsma and W. Beu-kelman, 1967.
97 S On the efficacy of two different roll-damping tanks. J. Bootsma and J. J. van den Bosch, 1967.
98 S Equation of motion coefficients for a pitching and heaving des-troyer model. W. E. Smith, 1967.
99S The manoeuvrability of ships on a straight course. ,J.. P. Hooft,
1967.
100 S Amidships forces and moments on a CB = 0.80 "Series 60"
model in waves from various directions. R. Wahab, 1967. 101 C Optimum conditions for blast cleaning of steel plate. Conclusion.
J. Remmelts, 1967.
102 M The axial stiffness of marine diesel engine crankshafts,. Part I. Comparison between the results of full scale measurements and
those of calculations according to published formulae. N. J.
Visser, 1967.
103 M The axial stiffness of marine diesel engine crankshafts. Part 11. Theory and results of scale model measurements and comparison with published formulae. C. A. M. van der Linden, 1967. 104 M Marine diesel engine exhaust noise. Part I. A mathematical model.
J. H. Janssen, 1967.
105 M Marine diesel engine exhaust noise. Part II. Scale models of
exhaust systems. J. Buiten and J. H. Janssen, 1968.
106 M Marine diesel engine exhaust noise. Part III. Exhaust sound,
criteria for bridge wings. J. H. Janssen en J. Buiten, 1967.
107 5 Ship vibration analysis by finite element technique. Part I.
General review and application to simple structures, statically
loaded. S. Hylarides, 1967.
108 M Marine refrigeration engineering. Part I. Testing of a
decentral-ised refrigerating installation. J. A. Knobbout and R. W. J.
Kouffeld, 1967.
109 S A comparative study on four different passive roll damping tanks. Part I. J. H. Vugts, 1968.
110,S Strain, stress and flexure of two corrugated and one plane
bulk-head subjected to a lateral, distributed load. H. E. Jaeger and
P. A. van Katwijk, 1968.
111 M Experimental evaluation of heat transfer in a dry-cargo ships'
tank, using thermal oil as a heat transfer medium. D. J. van der Heeden, 1968.
112 S The hydrodynamic coefficients for swaying, heaving and rolling cylinders in a free surface. J. H. Vugts, 1968.
113 M Marine refrigeration engineering. Part II. Some results of testing a decentralised marine refrigerating unit with R 502. J.. A. Knob-bout and C. B. Colenbrander, 1968.
114.S The steering of a ship during the stopping manoeuvre.. J. P.
Hooft, 1969.
115 S Cylinder motions in beam waves. J. H. Vugts, 1968.
116 M Torsional-axial vibrations of a ship's propulsion system. Part I. Comparative investigation of calculated and measured torsional-axial vibrations in the shafting of a dry cargo motorship.
C. A. M. van der Linden, H. H. 't Hart and E. R. Dolfin, 1968.
117 S A comparative study on four different passive roll damping
tanks. Part II. J. H. Vugts, 1969.
118 M Stern gear arrangement and electric power generation in ships propelled by controllable pitch propellers. C. Kapsenberg, 1968.
119 M Marine diese1 engine exhaust noise. Part IV. Transferdamping
data of 40 modelvariants of a compound resonator silencer.
J. Buiten, M. J. A. M. de Regt and W. P. H. Hanen, 1968.. 120 C Durability tests with prefabrication primers in use of steel plates.
A. M. van Londen and W. Mulder, 1970.
121 S Proposal for the testing of weld metal from the viewpoint of
brittle fracture initiation. W. P. van den Blink and J. J. W. Nib--bering, 1968.
122 M The corrosion behaviour of cunifer 10 alloys in seawaterpiping-, systems on board ship. Part I. W. J. J. Goetzee and F. J. Kievits,,
1968.
123 M Marine refrigeration engineering. Part III. Proposal for a
specifi-cation of a marine refrigerating unit and test procedures. J. A.
Knobbout and R. W. J. Kouffeld, 1968.
124 5 The design of U-tanks for roll damping of ships. 131,J van den
Bunt, 1969.
125 S A proposal on noise criteria for sea-going ships. J. Buiten, 1969'. 126 S A proposal for standardized measurements and annoyance rating
of simultaneous noise and vibration in ships. J. H. Janssen, 1969. 127 S The braking of large vessels II.H. E. Jaeger in collaboration with
M. Jourdain, 1969.
128 M Guide for the calculation of heating capacity and heating coils for double bottom fuel oil tanks in dry cargo ships. D. J. van der Heeden, 1969.
129 M Residual fuel treatment on board ship. Part III. A. Mooy, P. J. Brandenburg and G. G. van der Meulen, 1969.
130 M Marine diesel engine exhaust noise. Part V. Investiga jon of aJ double resonatorsilencer. J. Buiten, 1969.
131 S Model and full scale motions of a twin-hull vessel. H. F.. van
Sluijs, 1969.
132 M Torsional-axial vibrations of a ship's propulsion system. Part II, W. van Gent and S. Hylarides, 1969.
133 S A model study on the noise reduction effect of damping layers aboard ships. F. H. van To!, 1970.
134 M The corrosion behaviour of cunifer-10 alloys in
seawaterpiping-systems on board ship. Part II. P. J. Berg and R. G. de Lange, t.
1969.
135 S Boundary layer control on a ship's rudder: J1 H. IG. illerhagen,
1970.
136 S Observations on waves and ship's behaviotir made on board
of Dutch ships. M. F. van Sluijs and J. J. StiMman, 1971. 137 M Torsional-axial vibrations of a ship's propulsion system, Part HI.
C. A. M. van der Linden, 1969
138 S The manoeuvrability of ships, at low speed. J. P: Hooft and
M. W. C. Oosterveld, 1970.
139 S Prevention of noise and vibration annoyance aboard a sea-going
passenger and carferry equipped with diesel engines Part I. Line of thoughts and predictions. J. Buiten, J. H. ,Janssen,
H. F. Steenhoek and L. A. S. Hageman, 1971.
140 S Prevention of noise and vibration annoyance aboard a sea-going
passenger and carferry equipped with diesel engines. Part II. Measures applied and comparison of computed values with
measurements. J. Buiten, 1971.
141 S Resistance and propulsion of a high-speed single-screw cargo
liner design. .1. J. Muntjewerf, 1970.
142 S Optimal meteorological ship routeing. C. de Wit, 1970.,
143 S Hull vibrations of the cargo-liner "Kaudekerk". H. H. 't Hart,.
l9
144 S Critical consideration of present hull vibration analysis. S. Hyla-rides, 1970.
145 S Computation of the hydrodynamic coefficients, of oscillating
cylinders. B. de Jong, 1973.
146 M Marine refrigeration engineering. Part IV. A Comparative study on single and two stage compression. A. H. van der Tak, 1970. 147 M Fire detection in machinery spaces. P. J. Brandenburg, 11971. 148 S A reduced method for the calculation of the shear stiffness of a
ship hull. W. van Horssen, 1971.
149' M Maritime transportation of containerized cargo. Part II. Experi-mental- 'investigation concerning the carriage of green coffee from Colombia to Europe in sealed containers. J. A. Knobbout, 1971.
=
J.
