REPORT No. 121 S October 1968
NEDERLANDS SCHEEPSSTUDIECENTRUM TNO
NETHERLANDS SHIP RESEARCH CENTRE TNOSHIPBUILDING DEPARTMENT LEEGHWATERSTRAAT 5, DELFT
PROPOSAL FOR THE TESTING OF WELD METAL
FROM THE VIEWPOINT OF BRITTLE FRACTURE
INITIATION
(EEN VOORSTEL VOOR HET BEPALEN VAN DE WEERSTAND VAN GELASTE VERBINDINGEN TEGEN HET ONTSTAAN VAN BROSSE BREUKEN)
by
Jr W. P. VAN DEN BLINK Philips' Welding Electrodes Factory
and
Jr J. J. W. NIBBERING
Ship Structures Laboratory Delft Technological University
In het kader van de vraag naar optimale constructies wordt veel onderzoek verricht naar het sterktegedrag van scheepsconstruc-ties onder invloed van een realistisch optredende belasting.
In dit licht bezien is het eveneens van belang met meer zeker-heid dan thans mogelijk is de sterkte van lasverbindingen in deze constructies te kunnen bepalen. De in het algemeen optredende dynamische belastingen en het voorkomen van onvolkomen-heden in de las en overgangszone dienen mede in beschouwing te worden genomen.
De resultaten van daarop gericht onderzoek zijn in dit rapport
gegeven in de vorm van een voorstel voor een nieuwe
beproevings-methode van laswerk. Rekening is gehouden met de verkregen ervaring en overwegingen voor een praktische uitvoering in in-dustriele laboratoria. Vanzelfsprekend is de voorgestelde me-thode ook onderwerp geweest van diepgaand overleg in enige werkgroepen van het International Institute of Welding" (1.I.W.).
In principe komt de beproevingsmethode neer op het eisen van een bepaalde vervormbaarheid van het lasmetaal aan de voet van scherpe kerf. De afmetingen, een standaardkerf en de belasting worden gespccificeerd.
Gezien de belangrijke aspecten van het voorstel, mag een alge-mene aanvaarding van deze NIBLINK beproevingsmethode ten sterktste worden aanbevolen.
NEDERLANDS SCHEEPSSTUDIECENTRUM TNO
In the scope of the demand for optimum structures much research is performed into the strength behaviour of ship structures sub-jected to a reaslistic load.
In this respect it is also of importance to obtain more reliable information on the strength of welds in these structures. The generally prevailing dynamic loads and weld defects have to be taken account of.
The results of research in this direction are reportes in this publication in the form of a proposal for a new testing method of weldments. The method is based on present stage experience and practical application possibilities for industrial laboratories. Of course the proposed method has been subjected also to extensive discussions in some commissions of the International Institute of Welding (1.1.W.).
In principle the testing method comprises the requirement of a specified ductility of the weld metal at the root of a sharp notch. The dimensions, a standard notch and the loading have
been specified.
In connection with the important aspects of the proposal, a
general acceptance of this NIBLINK test may be strongly advocated.
NETHERLANDS SHIP RESEARCH CENTRE TNO
CONTENTS
page
Summary 7
1 Introduction 7
2 Description of test 7
3 Test piece and test procedure 8
4 Determination of critical C.O.D. values 11
5 Alternative procedure for the testing of type T test piece 12
6 Additional observations with regard to the operation of the test 13
7 Discussion of arguments leading to the proposed test 13
8 Summary of qualities of the proposed test method particularly with reference to the
Charpy-impact test 16 References 17 Appendix I I 8 Appendix II 19 . . . . . . . .
1 Introduction
The proposal presented in this paper is the result of a Netherlands investigation in the framework of a joint working programme of the International Institute of Welding (1.1.W.)-Working Group "Brittle Fracture Tests for Weldmetal" (W.G. 2912). One of the main
tasks of this working group, in which members of
commissions II, IX, X and XII are cooperating, is to device evaluation tests for weld metal from the view-point of brittle fracture danger. This task setting has
emanated from a wide-spread conviction that the
mechanical properties that determine the service be-haviour are in many cases not adequately reflected by the results of conventional small scale brittle
frac-ture tests.
As a result of its studies the working group has
con-cluded that for the majority of structures the
investi-gation of the brittle fracture initiation properties of the
weld are relatively more important than that of crack
arresting properties.
A further conclusion has been that an evaluation test for the weld metal should preferably involve a
parent metal/weld combination of the full plate
thick-ness to be applied.
It is generally agreed upon that brittle fracture initiation in a steel weldment normally involves the presence of a crack-like defect, and that, therefore, crack initiation tests should use test pieces containing crack-simulating sharp notches. Crack initiation thus is equivalent to "crack extension".
A matter of careful consideration has been how to
account for an unequality in yield value of parent metal and weld metal, which may be particularly
noticeable in mild steel weldments. This unequality results in different behaviour of the weld metal,
de-pending on the direction of the weld in a structure
relative to the main service stresses. Likewise it was
PROPOSAL FOR THE TESTING OF WELD METAL FROM
THE VIEWPOINT OF BRITTLE FRACTURE INITIATION
by
Jr. W. P. VAN DEN BLINK and
Jr. J. J. W. NIBBERING
Summary
A method of testing weldmetal for its sensitivity to brittle crack initiation is described. The method is based upon considerations derived from the present stage of experience and on considerations of feasibility by industrial laboratories.
Interested parties are invited to carry out tests on the basis of the proposal in order to investigate the practicability of the test and eventually to contribute to a collection of data necessary to improve testing requirements.
necessary to provide the possibility to test the weld in its two main directions, longitudinal and transverse.
A consideration in designing the test method has also been that the method should enable to provoke fractures in mild steel welds at temperatures not too far below environmental temperature without general
yielding of the weldment.
With the aim of ensuring a wide practicability of the test, its design has been chosen such that the test may be expected to be feasible for any interested in-dustrial laboratory.
The subdivision of the report is such that the first 6
sections mainly deal with the operation of the test, while in sections 7 and 8 the underlying philosophy is
given.
Important: as a result of the discussions on the proposal in Warsaw in 1968 some important modifications, par-ticularly with regard to the drop heights and the critical Crack Opening Diplacements (C.O.D.), have been
introduce& see section 2 and table I.
Moreover results for actual welds and information about the influence of variations in drop height and
notch sharpness can be found in the appendices.
2 Description of test
Essentially the method comprises the drop-weight loading by a series of consecutive blows of increasing
and defined height, and defined weight on a sharply
notched specimen containing the weld metal to be investigated.
Each subsequent blow results in an increased plastic
deformation in the notch tip region of the test piece. The amount of local plastic deformation before
frac-ture is used as a measure of the ductility. The usual type
of drop-weight test equipment, modified in some de-tails as will appear later, is suitable for the execution
of the test.
8
For evaluation purposes the test is carried out at different temperatures. Two types of welded test pieces may be used:
Type T containing a Transverse weld, notched in the centre line of the weld in the weld direction
(figure 2a).
Type L containing a Longitudinal weld, notched
per-pendicularly to the weld direction with the
notch extending to the centre of the weld (figure 2b).
When the test is used as an acceptance test, only one test temperature will normally be necessary. Three specimens are thought to be sufficient. As will be
ex-plained further, for acceptance testing type T has been
chosen as the standard specimen.
The critical amount of deformation to be required
is chosen as the total deformation undergone by a
test piece of unwelded plate material (type P), loaded to a calculated nominal stress equal to the minimum specified proof stress value in a static bend test (see
figure 1). This applies to non-stress-relieved structures; For stress-relieved structures half of that value is con-sidered to be sufficient, of course the specimens should
also be heat-treated.
The test method implies the necessity of measuring the deformation at the notch tip before or at fracture as advocated in particular by Wells. Several methods can be used, such as the measurement of the plastic zone size, of the lateral contraction at the notch root, of the total crack opening displacement or of the resi-dual crack opening displacement (C.O.D.).
The latter method mentioned has been chosen. Tests in the Delft Ship Structures Laboratory have
shown that when the deformations are small there is
an essential difference between the type of deformation
in notched bars in static and in drop-weight testing respectively. In a static test the residual contraction
in way of the notch is rather wide-spread and relatively
shallow. For the same amount of C.O.D. the
contrac-tion caused by drop-weight loading is much larger and occurs in a smaller region (pit-like). On account of this it has been decided that for drop-weight-specimens the
use of C.O.D. measurements should be preferred to
contraction measurements. Of course the only practical
possibility is measuring the C.O.D.'s after each blow is given (residual C.O.D.'s).
In the static test (figure 1) the total C.O.D. at yield point and not the residual C.O.D. after unloading
is used for the critical residual C.O.D. value in the drop-weight test. In this way the effect of residual welding stresses is also taken into account (see sect. 7). Consequently welds in steels of higher yield point will always have to meet a more severe notch deformation criterion which is reasonable with a view to the higher
325
Critical residual C.O.D. for drop weight test = Total C.O.D. at incipient yield in static test on base material.
Incipient yielding is defined to occur at the load at which the calculated nominal bending stress in the notched section is equal to the minimum specified Gy (kg/min')
proof stress value ---,P=116,.t t (mm) P (kg)
When stress relieving is effectuated by heat-treatment, the drop-weight specimens should be subjected to the same treatment before testing.
The critical C.O.D. then is half of the value indicated above.
Fig. 1 Determination of critical residual C.O.D. for modified drop-weight test with the aid of static bend test.
total deformation to which high strength steels are
subjected and with a view to the higher residual
stresses.
It will be clear that the results obtained with the proposed method will show a certain amount of scatter as a consequence of the inhomogeneity of weld metal. In principle for a pure initiation test, as the present
one, any specimen out of the prescribed number (3) should fulfill the requirement. On the other hand the test method as well as the criterion have the character of a compromise in order to be applicable to all kinds of practical situations. On account of this it is thought to be reasonable to allow that for acceptance-testing one out of three specimens shows a critical C.O.D.
slightly below the required one, provided that the
other two results are significantly better. This can be
met satisfactorily by prescribing that
C.O.D., x C.O.D., x C.O.D.3 > (C.O.D.cr;,)3
Assuming that for general application the residual
C.O.D. measuring is widely acceptable and easily
applicable (see figure 5), in the following sections reference will be made mainly to this method.
3 Test piece and test procedure
The test piece T and the test set-up are represented schematically in figure 2a. The test piece L is shown in
figure 2b. The V-butt weld is by way of example. The
test piece of the unwelded plate (P) has the same dimen-sions as the pieces T and L but machining of the surface in the notched region is not necessary except when
spe-cial measurements or observations are desired in this
region; the rolling direction of the plate should be Assumed stress
distribution .1
Hardened steel Rounded edges 3.25 !Hardened steel MOS Fig.. 3 Test-equipment 11 02 360 ipz 0.91 1.1 t (kg ,)1
L2
Detail of notch !region see also i.5 ) 2=
Rounded! edges Bridge pieceFig.. 2a teat piece 'T and the way of loading (proposed for acceptance testin,g)5
ii
See detail.
LI lt
Machined; flush with plate (both -sides/
Center of weld
U11(1 ((MIL_
2b Test piece t,
Fig. 2c Test piece ?'
Conditions for maximum thickness t 65 mm
When t> 65 mm the total height of the test piece should be f+2A/r mm., The bridge piece is omitted, the flat bottom part of the drop-weight should have a diameter of 75 mm and the span should be made 007t2.
r] 1;1 1. 4,-e --rmeszrart-A,, Fig. =
10
noted. As indicated in figure 2a, the test piece is sub,
jected to 4-point-bending in order to have pure bending of the notched part. The "bridge-piece" can best be
clamped to the test-specimen so that both are cooled at the same time (figure 3). The supports are flattened in order to limit the amount of energy lost by plastic deformation at the place of impact.
A rigid foundation of the supports is required' with
la view to the reproducibility of the test results (fi-.
sure 3).
The dimensions Of the' test piece are shown in figures
2a and 2b. As will be seen, the notch depth is made proportional to the square root of the plate thickness
and equal to 2.,/t, where t is the plate thickness in mm.
This relation has been chosen as a practical
compro-mise between a constant notch depth and a notch
depth proportional to the plate thickness, as will be
,explained in section 7.
The 0,2 mm notch can be made by hand sawing
with a saw-blade ground to the reguired thickness or a jeweler's saw.
The drop-hammer mass is related to the plate thick-ness and corresponds to a weight in kg equal to the
plate thickness in mm with a tolerance of + 10%. (See appendix I.c). In figure 4 a device is shown which allows quick and safe manipulating with the drop-hammer..
J.
Very reliable results (Adjusting of pin of extensometer easy(
Although any type of static displacement measure-ment is, in principle, suitable, those methods which allow a measurement without removing the test piece, from the anvil are preferable.
For general application a mechanical dial gauge extensometer of the type shown in figure 5a measuring,
the widening of a milled slot in the notch or a drilled
hole, can be used. Figure 5b shows an alternative
method which is more accurate but less easily to per-,
form.
Fig :,5' Possible methods of measuring C.O.D.
When testing at a series of temperatures, it is prac-tical to start with the lowest temperature envisaged. When the test piece is at the 'desired temperature, the test starts by measuring the initial distance for the
C.O.D.-measurement.
A first blow is then given from a height of 250 mm and after that the reference distance is measured. The
difference with the intial distance iwthe residual C.O.D.
Table I. Dropheight sequence
Cheap specime
AA
(Pin of extensometer adjusted from origi-nal circular section
to Indicated section This is difficult to do without impairing the accuracy,). -Figure 5a Small balls jpressed into the material ' Figure 5bi Hl. 112 H3 H4 11-5
H6. H7 H8 H9
in mm 250 300 350 4001 450 500 550 6001 700 etc. _____.;!II__IL
rig. 4 Magnetic safety clutch (when current is switched on pin in magnet goes down)
Beforeiubjecting a test piece to drop-weight loading
the notch has to be prepared for the measurement of
the residual C.Q.D.'
The second blow is given from a height of 300 m (see
table 1), and the C.O.D. is measured again, This is continued with the height 'increasing in steps of 50 mm under 600, mm height and in steps of 100 mm
height until fracture occurs. The C.O.D. measured at the last blow before fracturing is noted as the fracture
CO. D'. at the test temperature.
During the test the temperature of the test piece should be watched and corrected by cooling if necessa-ry. For evaluation testing, for instance when comparing
different weld metals, test pieces are tested the same
way at temperatures 10° and 20. ° higher respectively. At these higher temperatures the C.O.D.
measure-ments for the lower blow energies may be omitted as far as they did not fracture the piece at a lower
tem-perature.
In figure 6 the procedure is given diagrammatically. The temperature T is plotted at the abcis and the
drop-height H at the ordinate. The residual s are
plotted 'with -the temperature ordinates as a base::
Slow ric 8 .60 -if S5 50 5 3 lk 2E 1, so 70 40 35 30 25 # 0.20 r 0./5 .0.10 I0.05
Critical; C.0 0.4 see fig. 9 );
+.5° .810°'
Fig,7' Maximum C.O.D.% before fracture derived from results of figure 6
Next the fracture C.O.D..'s are plotted as a function of
'temperature (figure 7).
A direct comparison between different weld metals in the same steel is possible in a diagram as given in 'figure 7..
It will' be obvious that for steels in which cracks,
in the zones adjacent to the weld are a real possibility, the quality of these zones may be decisive for the quality of the weldment and may be determined in a similar way.
4. Determination of critical C.O.D. values
A static 'bend test is carried out at -the intended test temperature on a test piece of the steel (type P), using,
the same span and loading fixture as for the drop-weight tests. The total C.O.D. value at incipient yielding is determined. This is defined to occur at the load
value at which the "nominal maximum stress value", 'calculated by taking the remaining cross section in way of the notch as the cross section of a hypothetical bend test piece, equals the minimum specified proof stress value. For the dimensions of figure 2 and four point bending the load computed with a 1111W is:
P = flay'
t ay: kg/mm' (yield-stress)t':. mm (plate-thickness)
P: kg (load)
In figure 8 not only the CO. D. at incipient yielding but
ll C.O.D.'s measured from zero-load on are given for
tests at two different temperatures. The enormous
capacity for 'deformation of this material is obvious.
In thestatic test at 50°C the total C.O.D.. at fracture is more than twenty times as large as the residual
C.O.D. at fracture in a drop-weight test at a
tempera-ture 50° higher (figure 7).
In figure 8 one of the curves is obtained with a specimen containing a fatigue-crack. Even this one
behaved quit satisfactorily at 50.°C. Figure 8 shows
further that for loads below yield point the C.O.D. in
creases approximately linearly with the load. For loads between the one for which a calculated nominal' stress of yield point value exists and the one at which a
'plastic hinge is formed the C.O.D. progressively
in-creases with the load. At stiff higher loads a very rapid increase of C.O.D.. occurs.
Originally it has been considered to use. the residual
C.O.D.-value after unloading from the load at which ta plastic hinge is formed as a measure for the critical
drop-weight C.O.D. However this value cannot be
determined easily and accurately because it is very sensitive for small differences in the load applied (see figure 8)1.. Therefore the much more accurately to
determinetotalC.O.D..,at incipient yielding was chosen;
it is about equal to theresidual"plastic hinge" C.O.D.
(see figure 8) for the test conditions chosen.
11 ....-.
Ur
' 092 FractureI racture m. Or N .068 111050i
.
1103W 11.028 Atte and - Fructure 040 mm 10,03 10.045 10.033 Q.051 m 1.045 031 _. Fracture *au mm 0.02710.025 41022 , 0.073 0.012 0.016 0.00 o. 6.007 1 0.008 C.0 I 'surf 1 1 i 1 + 5 +10Test .temperature OD)
Semi-killed steel 37'
Isotherm Robertson arrest temp. +20°C Claarpy 3.5 kgm/cm2 (20 ft.lb) ± 3°C
5.2 kgm/cm2 (30 ft.lb) ± 7°C
7 kgm/cm2 (40 ft.lb) ± 9°C
50% Crystalline ±17°C
Fig. 6 Example of testing 'diagram for given mild steel .18 mmi 1st, 2,nd 3rd blow I. not mea-d. C.O.D.' 450 Fracture)( -10° -5°
4500 4000 3500 1.5 Cy 1 3000 2500 2200 "E 2000 4c 1500 5 1000 500 11 10 16 Ti7n-rn-57w 15 ,c
-;4
14 13 12 6 2 ViSual Calculated 01 0.2 C.O.D. (mm)-2 saw cu tPlastic hinge Incipient yielding minimum specified proofstress)C.O.D. at incipient yield to be used as critical value for residual
C.O.D in drop weight test.
Residual can.after unloading
from Load at which plastic hinge
is formed.
It is obvious that the establishment of critical C.O.D. values is a procedure that can be carried out separately from any weld metal testing.
It is to be expected that the C.O.D.'s at incipient
yielding are un-ambiguously defined by yield point and plate thickness. After some experience has been gained there is no need for further static bend testing.
Important: C.O.D. measurements are carried out in
various ways in different laboratories. The
configu-ration of the notch might be made different from what is given in this report. Also the place where the C.O.D. measurements are taken might be chosen different
from what is proposed.
All these variations are acceptable provided that in
the static bend test exactly the same C.O.D. measuring is applied as in the drop-weigth testing. In this way the
influence of measuring techniques on the final result
can be eliminated.
Alternative procedure for the testing of type T test piece
The procedure as described above is based on the
assumption that welds in a structure are subjected to the same amount of strain as the plate material,
irres-0.3
Semi-killed steel 37
Isotherm Robertson arrest temp. +20 °C Charpy 3.5 kgm/cm2 (20 ft.lb) 3°C
5.2 kgm/cm2 (30 ft.lb) + 7°C
7 kgm(cm2 (40 ft.lb) 9°C
50% Crystalline +17°C
Fig. 8 C.O.D. values obtained with static bend-tets at 5°C and 50 C for given mild steel
111
No fracture at C.0.D.055mm
.c No fracture atcovs=1,5mm
04 0.5 0.6
pective of the yield value ratio of steel and weld metal.
For longitudinal welds, with the main service stress
in the direction of welding, the validity of this assump-tion was confirmed by tensile testing of notched plates [6]. In this case the deformation of the plate is imposed on the weld, regardless of its yield value. For the L-test
piece, therefore, there is no doubt that equal perfor-mance of different weld metals means equal
defor-mation of the test pieces, i.e. equal C.O.D.
For welds in the transverse direction, however,
there are cases in which the deformation of the weld
is not governed by the overall deformation of the weldment but by the ratio of stress to weldmetal yield
value. It is obvious that the deformations at notches
in an uninterrupted weld, loaded perpendicularly to its
direction by a uni-axial stress field will be governed only by the magnitude of the applied. stress If the
yield value of the weld is lower than that of the plate, a notch in the weld will start to deform plastically at a lower load than the plate and vice versa.
To take account of these conditions, an alternative mode of evaluation is possible, which will be denoted
T,.
To start with, the procedure involves the
drop-weight loading of unwelded test pieces with the aim 12
Calculated Plastic
of finding the drop-height at which such a test piece
shows a C.O.D. equal to the critical value. The welded test piece T is then loaded by the same number of
stepwise increasing drop-heights, resulting in
corre-sponding nominal stress fields as in the unwelded P-spe-cimen. If the test piece does not fracture it is considered to have fulfilled the T' test requirement.
It will be obvious that for mild steel weldments the T' procedure will normally be milder than the regular
T-test, because the higher yield point of the weld metal will give rise to a smaller deformation in the weld metal as compared with the base material for the
same drop-height. For alloy steels, however, the T' procedure may be more severe than the regular test, if the yield value of the weld metal happens to be lower than that of the steel.
6 Additional observations with regard to the operations of the test
From the foregoing sections the impression may be
gained that multiple drop-weight testing involves more
work than the determination of for instance a Charpy
V temperature curve.
This may be true in the present stage. It is to be expected however, that after some experience has been gathered, the procedure can be very much simplified.
For instance the static bend tests will be very soon
no longer necessary when for a few yield value classes of steel and plate thicknesses (and perhaps methods of
C.O.D. measuring) reliable data are obtained to make "master-charts" for the critical C.O.D.
The choice between the L and T-type of test piece can be easily made if there is no doubt about the way
of loading of the weld in the structure.
A difference between the result of L- and 1-testing
may be expected from the crystallisation texture of the
weld, which may tend to favour fracture in the weld
direction.
Apart from the relative severity of L- versus 1-testing, it is worthwhile to consider what relative weight should be attached to the result of each test with a view
to the type of defects to be expected in actual welds. If defects transverse to the weld are expected to be virtually non-existent, it may be justified to omit the
L-testing altogether.
It may be taken for granted that by far most of the
defects resulting from the welding procedure itself
tend to be in the direction of welding, so that in gen-eral the T-type of test cannot be omitted.
For this reason testing of the T-type is proposed as the
standard procedure. This simplifies the process and has
the additional advantage that little material is con-sumed.
The number of blows necessary to give a certain
C.O.D. in a given test piece may depend on the rigidity of the foundation of the supports. But the final C.O.D. to fracture will hardly be influenced by small variations in rigidity. A satisfactory set-up is shown in figure 3.
It sometime happens during a test that at a certain
blow the residual C.O.D. increases far more than during the preceding steps. This can be due to the
for-mation of an internal crack ("tunnel-crack-) not
visible from the outside.
When this is suspected to have occurred it is re-commended to insert ink in the notch, after which the
specimen can be fractured in order to inspect the frac-ture-surface.
In the foregoing sections the testing of the parent
steel is included mainly to have a basis for relating the applied stress level to the resulting notch deformations.
It has not been the intention to compare steel
perfor-mance versus weld perforperfor-mance with the aim to predict the actual performance in a structure.
For such a comparison not the parent metal but the most critical zones surrounding the weld should be investigated preferably in the same way as proposed
for the weldmetal.
7 Discussion of arguments leading to the proposed test a. In the study of brittle fracture phenomena it is
necessary to differentiate between the propagation
aspects, which concern the global weldability proper-ties, and the initiation aspects, for which local
con-ditions are to be considered. It needs hardly to be
said that the most critical regions for crack initiation are the welds and the zones directly adjacent to it.
Even for those structures, for which the design and
choice of material is based upon considerations of crack arresting, the safety of operation is improved by measures that limit the danger of crack initiation.
For an important category of weldments the crack
propagating behaviour of the weld metal is considered
not to be of primary interest for the safety of the structure as a whole. Both the statistical evidence from brittle fracture cases in service [1] and the experimen-tal evidence from tests on large welded plates [2], [3].
[4] indicate that under certain conditions brittle frac-tures tend to avoid the weld proper. Although it is not
possible to outline quantitatively the physical
con-ditions which determine the fracture path, it is still
possible to distinguish in technical terms specific cases. The following circumscription is considered to cover
those cases for which the existing evidence predicts
that a brittle fracture, even if initiated in or in the vicin-ity of a weld, will not follow the weld proper:
Butt welds, welded through the complete plate thick-ness by means of conventional multi-layer processes, in nonstress relieved structures of ferritic steels with
minimum specified proof-stress values not greater than 40 kg/mm2.
For these structures the safety of operation is best
guarded by requiring crack arresting properties of the
steel and freedom from crack initiation in the weld and its environment.
For structures outside this category, the crack
arrest-ing properties of the weld metal need consideration,
but it will be obvious that the prevention of crack
initiation remains a matter of importance. From the
foregoing it may be clear that the I.I.W. Working
Group "Brittle Fracture Tests for Weld Metal" has
concentrated first its activities on the initiation aspects
of brittle fractures.
b. In the introduction to this paper it has been stated
that there is a general conviction that conventional
notch ductility testing, particularly Charpy-impact
testing, does not adequately reflect the expected service behaviour of a weld. This does not mean that the
con-ventional testing results have no relation to the duc-tility properties of the metal but that there is no reli-able method to relate the results of conventional
me-thods with that to be expected in actual structures. This uncertainty is basically caused by the difficulty to
relate quantitatively the differences in behaviour of a ductile metal under conditions of different geometry
and loading speed.
A practical difficulty encountered in Charpy testing
is that, in the transition range, a very large scatter in impact values is often found and that there is no
uni-formity of opinion about the significance of this scatter from the viewpoint of material evaluation.
Uncer-tainty also exists with regard to the different values
found in different regions of a multi-layer weld in heavy
Fig. 9 Specimen with lack of penetration
plates. Finally it is quite impossible to obtain some idea about the influence of weld faults, like lack of
penetration (figure 9), porosity etc. on the resistance to
brittle fracture with the aid of small size specimens.
These considerations have become more and more
im-portant in the last ten years when the welding of thick plates has become common practice.
Given this situation, a logical approach is to
in-vestigate the metal under conditions which are more closely related to that under service conditions. From
a technical viewpoint the testing of structure simulating
elements is the best solution, but this method is not suited for general application.
So there seems to be room for test methods on
smaller test pieces of which the results can be applied
with more confidence to structural behaviour than
that of conventional notch ductility tests.
A test for crack initiation on relatively wide plates
has been proposed and investigated by Ikeda et al. [5]. In this deep-notch test, the loading is purely static,
which could explain the reported rather low initiation temperatures. The test is not suited and not meant for general application because it involves the use of expensive testing equipment.
On the other hand in a smaller test piece, such as
proposed in the present paper, not all conditions
acting in structures, such as the effect of very deep notches and of residual stress fields, can be included.
Therefore there remains room for tests on larger assemblies to account for such influences.
For most cases however the test procedure proposed will satisfactorily simulate the actual conditions to
which a weld is submitted in a structure.
One of the crucial questions has been to what extent
weld metal will be deformed in an actual structure if notches are present. To get an answer to this problem
an investigation was carried out by Nibbering [6] with
the aid of extensive measurements on a tensile loaded
plate, welded in the tensile direction and provided with several notches, comprising different zones in the weldment.
This test showed very clearly that the notch in the weld metal is plastically deformed to the same extent
as the virgin plate material. It is obvious that transverse welds will mostly not behave in the same way, but it is believed to be a safe procedure to assume that in many cases a transverse weld as well will have to deform along
with the plate. This may be the case for instance in
transverse welds in flanges in composite beams and at weld crossings.
Nevertheless the test procedure T' leaves room for
those cases in which the design is such that transverse,
uninterrupted welds are loaded by a linear stress field
over their entire length (see section 5). In that case the
14
weld should be subjected to a stress criterion rather than a deformation criterion. In case the yield value of the weld is lower than that of the plate steel the stress criterion (i.e. the T' procedure) is the more severe one and should be applied.
c. A dynamic test has been chosen in the first place to provoke fractures at temperatures not too far below room temperatures and this making the test less cumbrous for general application. This choice was supported by results of tests on specimens containing fatigue cracks reported by Nibbering et al. [7]. Static or dynamic testing proved to make a difference of more than 50 C in transition behaviour. A second reason for testing by an impact load was that for many struc-tures the occurrence of shock is a real danger; shocks may happen either by accident or as result of small
local brittle fractures which may develop during
fatigue loading, when the fatigue crack travels through the various zones of a weldment (figure 10). A third reason is that the test-equipment turns out to be simple and in expensive.
Fig. 10 Fracture-surface of H.A.Z. (Heat-affected zone) of electro gas welded plate (result of fatigue loading at
20°C)
The intermittent impact loading has been chosen partly in order to get more information out of one test piece than is possible by loading by a single stroke:
the latter is essentially a go no go test.
A second reason to chose intermittent loading by increased drop heights is that it is believed that in this way real loading conditions are more closely
ap-proached than by single blow loading. This opinion is based on the consideration that the blows that do not fracture the test piece, will cause a condition of defor-mation and stress around the notch tip, which will be less different from the condition that results from static loading than in a one-blow test. The energy from a next blow will mainly be consumed in the elastic deformation of the test piece and only a final fraction of the energy will further deform the notch tip region plastically and eventually lead to fracture. Thus the condition of fracture may be expected to be rather similar to what happens if a tensile loaded specimen is fractured by a blow of low energy.
In fact a compromise is obtained between pure
static loading and the high speed shock loading in-herent to normal impact testing. Important is that in this way the phenomenon of strain hardening at the notch tip is retained.
The drop-height steps and the hammer weight,
which determine largely the speed of loading, have been chosen on the basis of what is thought to be realistic and are connected with the experience obtained
so far [7], [8].
The initial height is 250 mm for all steels.
Conse-quently higher strength weld metals will suffer a
greater number of blows until the critical C.O.D. value is attained than lower strength weld metals. This might seem to be of advantage for the higher strength metals because in general the greater the number of
blows the larger the C.O.D. before fracture, (see
appendix lb). But it should be realized that at the
moment of fracture higher strength metals are sub-jected to a higher speed of loading than lower strength metals because the final drop-height is larger. More-over the critical C.O.D. prescribed for higher strength steels is larger than for lower strength steels being prac-tically proportional to yield point. The main reason is that in structures the total C.O.D.' s at notches are higher for high strength steels as compared to mild steel. For non stress relieved structures the critical C.O.D. should of course be higher than in stress relieved structures. Static loading to yield point of a non-stress relieved structure will after unloading generally result in residual plastic deformations of a magnitude equal to o,/ E. For the case of moderate dynamic loading the presence of residual stresses will have a similar effect. That is why the critical C.O.D. has been defined as it is. For stress-relieved structures half of that value is con-sidered to be sufficient.
The form of the notch is that of a machined slot
with a width of 0.2 mm at the bottom. A natural
crack, for instance a fatigue crack, is obviously at-tractive, but has the disadvantage of being more diffi-cult to make in a reproducible way. The speed of
formation at the bottom of a saw-cut is lower than of a fatigue-crack, which meets once again the wish to
avoid a too extreme dynamic character of the test.
An-other disadvantage of fatigue-cracks for the type of
test used is that, after one or more blows, such a crack
is relatively much more blunted than a saw-cut notch,
resulting in an uncontrolled shift in the severity of the
test. In appendix Id some information is given about the difference in behaviour of specimens containing saw-cuts or fatigue-cracks. The difference is partly a result of the sharpness of the crack and partly of the deterioration of the material at the tip of the crack.
On the whole it amounts to about 25 'C difference
in critical temperature.
The depth of the notch has been chosen in propor-tion to the root of the plate thickness. This is a
com-promise between a constant depth and proportionality
to the thickness. A constant depth is attractive from the viewpoint of interpretation of the results. With a view to the technological character of the test,
how-ever, an increase of the notch depth with the plate
thickness seemed appropriate in order to account for the fact that real defects normally will also tend to in-crease in length with plate thickness, if only by the decreased probability of detection. The square root was chosen because it was assumed that a linear rela-tion would exaggerate the effect of plate thickness on
defect size and also to keep the test piece within
wield-able dimensions for the greater plate thicknesses.
The overall dimensions have mainly been chosen in connection with the existing experience. The height of the remaining section in way of the notch, the ligament,
has been made constant, 65 mm, because otherwise
the testing conditions would have to be varied in a
more complicated way. As it is now, only the
drop-weight has to be varied in proportion to the plate
thickness to induce the same "nominal- stress field at
the same drop height in different plate thicknesses.
The deformation and stress at the notch tip will, of
course, vary with the plate thickness, butt this is exactly a cause of the size effect to be accounted for. A second reason to keep the ligament constant was to have a constant gradient of the nominal bending stress in way of the notch for a given yield-value class.
The 65 mm for the ligament is believed to be
suffi-ciently large, so that a plastic strain field in the tip
region is not strongly disturbed by the vicinity of the neutral axis and does not differ too much from
nor-mal tensile conditions.
8 Summary of qualities of the proposed test method
particularly with reference to the Charpy-impact test
a. Initiation characteristics are clearly separated from propagation characteristics.
Quality has been defined in terms of directly
measured local ductility (C.O.D.) instead of in terms of a complex figure like specific energy.
Full plate thickness, so thickness-effect has been included and the inhomogeneous character of the weld has been taken into account.
Notch size and acuity conform rather to realistic
cracks.
Strain hardening as occurs in static loading, has
been maintained by applying progressively
in-creasing drop heights.
Realistic compromise between static and
conven-tional impact tests. (Strain rate is restricted by
using low drop heights - stepwise increased - and
a saw-cut notch instead of a natural crack).
Possibility of comparing weld and heat-affected
zone (H.A.Z.)
Influence of weld defects, for instance lack of pene-tration, porosity etc. on resistance to brittle fracture can be estimated.
Apart from the above it is attractive that
crack-arresting properties of the weld can simply be estimated
with same specimens by applying large drop heights [7], [9].
Objections to the proposed method can easily be found of course as always with compromises. The
choice of loading speed is arbitrary; the derivation of
the critical C.O.D. value from a pure static test is
theoretically not wholly justified; the notch is not a natural crack. However eventually the test can easily be adjusted to meet such objections if required. It is suggested that in order to limit the number of para-meters and to maintain the possibility to compare the results without the need of applying confusing
cor-rections only drop height, drop weight and critical C.O.D. value should be varied if necessary.
Some results with respect to the first two variables are given in appendix I. To show the outcome of the
method when applied to welded plates, results for
submerged arc-, electrogas- and automatic CO3-welds are collected in Appendix II. Comparisons with results obtained with Charpy-V-notch specimens can be made.
It is hoped that the interested parties, notably those represented in the I.I.W. Commissions If, IX, X and
XI I will be willing to contribute to obtain experimental
data by carrying out testing programmes on the basis of the proposal presented in this paper.
One of the main reasons for giving the proposed
testing parameters in rather great detail is to ensure the
possibility to compare testing results from different
sources in a cooperative investigation. 16
I
-References
I. Aupica, A., Progress reports of Working Group _Brittle Fracture in Service". IIW-doc. X-387-66, X-424-67 and Welding in the World, 1965, pp. 58-67.
KIHARA, H., Recent Studies in Japan on Brittle Fracture of Welded Steel Structure under Low Applied Stress Level. Japan Institute of Welding, 11W-doe. X-291-61.
DECHAENE, R. and J. SEBILLE, Euratom Colloquium on Brittle
Fracture. Proceedings, pp. 445-478.
SELANDER, L.. and L. TODELL, Brittle Fracture Propagation
in Welded Joints. Institute of Welding Technology,
Stock-holm. 11W-doe. IX-573-68/X-567-68.
IKEDA, K., Y. AKITA and H. K !NARA, The Deep Notch Test
and Brittle Fracture Initiation. 11W-doe. X-404-67.
NIBBERING, J. J. W., Plastic deformations at notches in welds
of mild steel plates. S.S.L. rep. 129, (11W-doe. 29/2-/07
1968).
NIBBERING, J. J. W., J. VAN LINT and R. T. VAN LEEMEN,
Brittle fracture of full-scale structures damaged by fatigue. Neth. Ship Researchcentr. Report no. 85 S. 1966.11W-doe. X-374-66 and I.S.P. Nov. 1966.
VAN DEN BLINK, W. P., A crack extension test for weld metal. 11W-doe. IX-527-67/X-453-67.
MARQUET, F., Side bend test procedure. Steel times, Nov. 19, 1965. 17 4, 5L, 6.
1
18
APPENDIX I
a. Influence of drop-height of first blow on test results
C.O.D
C.O.D.
Average of 2 "2 specimens
c. Influence of variations in drop-weight on test results
4 0.033 0.025 2 0.015 Fracture 500
i
1 0.010 : Blow COD. (mm) . nn Test temperature: -5 °C Average of 2a2 specimensConclusions:
Influence is moderate, (about
10% of critical C.O.D. for a difference of 50 mm in ini-tial height)
b. Influence of step-magnitude on test results
Test temperature 0 °C Average of 2 x 2 specimens
C.O.D.
Conclusions:
Influence is distinct, but when the results are com-pared with those of figure 6 can be seen that in
"transi-tion" temperature only a difference of about 5°C is obtained
d. Influence of sharpness of notch on test-results Comparison between saw-cut notch and fatigue-crack From the results d. to j. given in the following table it appears that the critical temperature for specimens containing fatigue-cracks with a length greater than 1 mm is about 30°C. (The critical C.O.D. was 0.04 mm; see figure 8). For specimens con-taining saw-cuts the critical temperature was about 0°C.
It is important to know if this large difference is only due to the difference in sharpness between saw-cuts and cracks or if it is partly due to the deterioration of the material caused by the fatigue-loading.
The results for the specimens a, b and c, containing very small fatigue-cracks (0.5 mm) suggest that only 10°C of the total of
30°C were caused by the mentioned deterioration.
N Length of crack (mm) Temp. (0 C) Height of first step Number of blows C.O.D. before frac-ture (mm) Observations a. 50,000 0.5 + 5 25 6 0.035 l length of b. 50,000 0.4 +10 25 9 0.240 fatigue-cracks c. 96,000 0.5 +20 30 5 0.360 J 0.5 mm d. 71,000 2.5 0 30 1 o e. 50,000 1 0 25 5 0.022 f. 70,000 2.2 +10 30 1 0 length of g. 98,000 1.5 +10 15 3 0.010 fatique-cracks h. 126.000, i. 101,500 3.5 1.8 +20 +30 15 15 6 6 0.037 0.030 I mm j. 110,000 +2.5 +40 15 6 0 80 500 /--Fracture : 550 Blow c.o.D.(mm) 500 3-6 0.041 0.054 450 5 0 045 400 2-4 0.020 0.033 350 3 0.025 300 1 2 0.007 0.016 250 1 0.008 200 ,...-Fracture 550 nlow no C.O.D. (mm) 500 0.049 0.054 450 --5 0.038 0.045 400 -4 0.027 0.033 350 -3 0.017 0.025 300 -2 0.009 0.016 250 1 0.008 200 2 Test temperature 0 °C
I.
5 0.040 W exi--4.x 0.031,
3 0.027 2r
0.012 0. --1 0.007 1000 900 800 700 600 Fracture Conclusions:Results conform if plotted Blow Carl(mm) on the basis of energy -5 0.046 (weight x height). Difference in fracture C.O.D. is not alarming (10%) 450 400 350 300 250 200 100
Summary of tests on welded plates (Dropweight 22 kg)
a Plate material:
(Semikilled, as rolled)
thickness 25 mm; yield point 25 kg/mm2
tensile strength 45 kg/mm2 Icritical C.O.D. 0.055 mm critical temp.I- 10°C (Charpy 3.5 kgm/cm 2 : I - 7 CI) 50% cryst. + 7°C v. d. Veen +8°C)
b Submerged arc weld: L-test
Symm. double V, 90', two passes, resp. 800 and
1000 Amp.
Weld speed 33 cm/min, wire 5 mm, cry = 35,8 kg/mm2, = 48,5 kg/mm2
Temp. C.O.D. mm
APPENDIX II
(Charpy 3.5 kgm/cm2: I 20°C )
d Electrogas weld :
T-test; gap 14 mm; one pass (enclosed CO2) solid
wire 1.6 mm; 450 Amp; 5 cm/min
Temp. C.O.D. mm Temp. C.O.D. mm
10°C
0 0°C 0.090 0°C 0.035 Results: Temp. C.O.D. mm20°C
0.01 10°C 0.045 10°C 0.070 0°C 0.280 1930°C
0.160 critical temp. I < 30°C I 20°C 0.150 (Charpy 3.5 kgm/cm2:f Electrogas weld (enclosed CO2): 1-test
Composite elestrode 2.4 mm; 450 Amp; 5 cm/min
Temp. C.O.D. mm
10°C 0.11
critical temp. I 10°C I10°C 0.165
(Charpy 3.5 kgm/cm2 : I + 25 °C ! I)10°C 0.050
(-10°C
0.045) fracture at copper inclusion0°C 0.400 next to notch)
0°C 0.140
0°C 0.190
e Automatic CO, weld :
L-test; symm X; 60'; hor; root pass: basic electr. Composite electrode (basic core) 2.4 mm
450 Amp; 45 cm/min; 8 passes Temp. C.O.D. mm
0°C 0.19 critical temp. 15°C I (estim.)
5°C 0.09 (Charpy 3.5 kgm/cm2: I -10°C
10°C 0.065
c Submerged arc weld:
T-test; with lack of penetration (see figure 9)
critical temp. :I (estim.)
40°C 0.041 critical temp. :I 18°C I 30°C 0.055 (Charpy 3.5 kgm/cm21 20 °C
20°C
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IS The determination of the natural frequencies of ship vibrations (Dutch). H. E. Jaeger, 1950.
3 S Practical possibilities of constructional applications of aluminium alloys to ship construction. H. E. Jaeger, 1951.
4 S Corrugation of bottom shell plating in ships with all-welded or partially welded bottoms (Dutch). H. E. Jaeger and H. A.
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5 S Standard-recommendations for measured mile and endurance trials of sea-going ships (Dutch). J. W. Bonebakker, W. J. Muller and E. J. Diehl, 1952.
6 S Some tests on stayed and unstayed masts and a comparison of experimental results and calculated stresses (Dutch). A. Verduin and B. Burghgraef, 1952.
7 M Cylinder wear in marine diesel engines (Dutch). H. Visser, 1952. 8 M Analysis and testing of lubricating oils (Dutch). R. N. M. A.
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10 S On collecting ship service performance data and their analysis.
J. W. Bonebakker, 1953.
11 M The use of three-phase current for auxiliary purposes (Dutch). J. C. G. van Wijk, 1953.
12IM Noise and noise abatement in marine engine rooms (Dutch). Technisch-Physische Dienst TNO-TH, 1953.
13 M Investigation of cylinder wear in diesel engines by means of labo-ratory machines (Dutch). H. Visser, 1954.
14 M The purification of heavy fuel oil for diesel engines (Dutch). A. Bremer, 1953.
15 S Investigations of the stress distribution in corrugated bulkheads with vertical troughs. H. E. Jaeger, B. Burghgraef and I. van der
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16 M Analysis and testing of lubricating oils II (Dutch). R. N. M. A. Malotaux and J. B. Zabel, 1956.
17 M The application of new physical methods in the examination of
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18 M Considerations on the application of three phase current on board ships for auxiliary purposes especially with regard to fault pro-tection, with a survey of winch drives recently applied on board of these ships and their influence on the generating capacity (Dutch). J. C. G. van Wijk, 1957.
19 M Crankcase explosions (Dutch). J. H. Minkhorst, 1957.
20 S An analysis of the application of aluminium alloys in ships' structures. Suggestions about the riveting between steel and aluminium alloy ships' structures. H. E. Jaeger, 1955.
21 S On stress calculations in helicoidal shells and propeller blades. J. W. Cohen, 1955.
22 S Some notes on the calculation of pitching and heaving in longi-tudinal waves. J. Gerritsma, 1955.
23 S Second series of stability experiments on models of lifeboats. B.
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24 M Outside corrosion of and slagformation on tubes in oil-fired boilers (Dutch). W. J. Taat, 1957.
25 S Experimental determination of damping, added mass and added
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27 S Initial metacentric height of small seagoing ships and the
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28 M Influence of piston temperature on piston fouling and pistonring wear in diesel engines using residual fuels. H. Visser, 1959. 29 M The influence of hysteresis on the value of the modulus of
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30 S An experimental analysis of shipmotions in longitudinal regular waves. J. Gerritsma, 1958.
31 M Model tests concerning damping coefficient and the increase in the moment of inertia due to entrained water of ship's propellers. N. J. Visser, 1960.
32 S The effect of a keel on the rolling characteristics of a ship.
J. Gerritsma, 1959.
33 M The application of new physical methods in the examination of lubricating oils (Contin. of report 17 M). R. N. M. A. Malotaux and F. van Zeggeren, 1960.
34 S Acoustical principles in ship design. J. H. Janssen, 1959.
35 S Shipmotions in longitudinal waves. J. Gerritsma, 1960.
36 S Experimental determination of bending moments for three
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37 M Propeller excited vibratory forces in the shaft of a single screw tanker. J. D. van Manen and R. Wereldsma, 1960.
38 S Beamknees and other bracketed connections. H. E. Jaeger and J. J. W. Nibbering, 1961.
39 M Crankshaft coupled free torsional-axial vibrations of a ship's propulsion system. D. van Dort and N. J. Visser, 1963.
40 S On the longitudinal reduction factor for the added mass of vi-brating ships with rectangular cross-section. W. P. A. Joosen and J. A. Sparenberg, 1961.
41 S Stresses in flat propeller blade models determined by the moire-method. F. K. Ligtenberg, 1962.
42 S Application of modern digital computers in naval-architecture. H. J. Zunderdorp, 1962.
43 C Raft trials and ships' trials with some underwater paint systems. P. de Wolf and A. M. van Londen, 1962.
44 S Some acoustical properties of ships with respect to noise control. Part. I. J. H. Janssen, 1962.
45 S Some acoustical properties of ships with respect to noise control Part II. J. H. Janssen, 1962.
46 C An investigation into the influence of the method of application on the behaviour of anti-corrosive paint systems in seawater. A. M. van Londen, 1962.
47 C Results of an inquiry into the condition of ships' hulls in relation to fouling and corrosion. H. C. Ekama, A. M. van Londen and P. de Wolf, 1962.
48 C Investigations into the use of the wheel-abrator for removing rust and millscale from shipbuilding steel (Dutch). Interim report. J. Remmelts and L. D. B. van den Burg, 1962.
49 S Distribution of damping and added mass along the length of a shipmodel. J. Gerritsma and W. Beukelman, 1963.
50 S The influence of a bulbous bow on the motions and the
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51 M Stress measurements on a propeller blade of a 42,000 ton tanker on full scale. R. Wereldsma, 1964.
52 C Comparative investigations on the surface preparation of ship-building steel by using wheel-abrators and the application of
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53 S The braking of large vessels. H. E. Jaeger, 1963.
54 C A study of ship bottom paints in particular pertaining to the behaviour and action of anti-fouling paints A. M. van Londen,
1963.
55 S Fatigue of ship structures. J. J. W. Nibbering, 1963.
56 C The possibilities of exposure of anti-fouling paints in Curacao, Dutch Lesser Antilles, P. de Wolf and M. Meuter-Schriel, 1963. 57 M Determination of the dynamic properties and propeller excited vibrations of a special ship stern arrangement. R. Wereldsma,
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58 S Numerical calculation of vertical hull vibrations of ships by discretizing the vibration system. J. de Vries, 1964.
59 M Controllable pitch propellers, their suitability and economy for large sea-going ships propelled by conventional, directly coupled engines. C. Kapsenberg, 1964.
60 S Natural frequencies of free vertical ship vibrations. C. B.
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61 S The distribution of the hydrodynamic forces on a heaving and
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62 C The mode of action of anti-fouling paints: Interaction between anti-fouling paints and sea water. A. M. van Londen, 1964. 63 M Corrosion in exhaust driven turbochargers on marine diesel
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64 C Barnacle fouling on aged anti-fouling paints ; a survey of pertinent literature and some recent observations. P. de Wolf, 1964.
65 S The lateral damping and added mass of a horizontally oscillating shipmodel. G. van Leeuwen, 1964.
66 S Investigations into the strenght of ships' derricks. Part I. F. X. P
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67 S Heat-transfer in cargotanks of a 50,000 DWT tanker. D. J. van der Heeden and L. L. Mulder, 1965.
68 M Guide to the application of Method for calculation of cylinder liner temperatures in diesel engines. H. W. van Tijen, 1965. 69 M Stress measurements on a propeller model for a 42,000 DWT
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70 M Experiments on vibrating propeller models. R. Wereldsma, 1965.
71 S Research on bulbous bow ships. Part II. A. Still water perfor-mance of a 24,000 DWT bulkcarrier with a large bulbous bow.
W. P. A. van Lammeren and J. J. Muntjewerf, 1965. =
