NEDERLANDS SCHEEPSSTUDIECENTRUM TNO
NETHERLANDS SHIP RESEARCH CENTRE TNO
SHIPBUILDING DEPARTMENT LEEGHWATERSTRAAT 5, DELFT
*
PROPOSAL FOR THE TESTING OF WELD METAL
FROM THE VIEWPOINT OF BRITTLE FRACTURE
INITIATION
(EEN VOORSTEL VC)OR HET BEPALEN VAN DE WEERSTAND VAN GELASTE VERBINDINGEN TEGEN HET ONTSTAAN VAN BROSSE BREUKEN)
by
Ir W. P. VAN DEN BLINK Philips' Welding Electrodes Factory
and
Ir J. J. W. NIBBERING
Ship Structures Laboratory Deift 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 bet algemeen optredende
dynamische belastingen en het voorkomen van
onvolkomen-beden 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-dustriële laboratoria. Vanzelfsprekend is de voorgestelde
me-thode ook onderwerp geweest van diepgaand overleg in enige
werkgroepen van het "International Institute of Welding"
(I.I.W.).
In principe komt de beproevingsmethode neer op het eisen van een bepaalde vervormbaarheid 'ian het lasmetaal aan de voet van scherpe kerf. De afmetingen, een standaardkerf en
de belasting worden gespLcificeerd.
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 (1.1W.).
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.
CONTENTS
page
Summary 7
i 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 1 2
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 18
PROPOSAL FOR THE TESTING OF WELD METAL FROM
THE VIEWPOINT OF BRITTLE FRACTURE INITIATION
by
Ir. W. P. VAN DEN BLINK and
Ir. J. J. W. NIBBERING
Summary
A method of testing weidmetal 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.
i
IntroductionThe 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 (I.I.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, TX, 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.
lt 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
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.
hnportant: 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 introduced; 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
Assumed stress distribution
Critical residual COD. for drop weight test = Total C.O.D. at
incipient yield in static test on base material.
Incipient yielding isdefined to occur at the load at which the
calculated nominal bending stress in the notched section is
equal to the minimum specified r (kg/mm2)
proof stress value -*P = I li.,.t t(mm)P(kg)
When stress relievingiseffectuated by heat-treatment, the
drop-weight specimens should be subjected to the same treatment
before testing.
The critical CO. 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.1 x C.O.D.2 x C.O.D.3 > (C.O.D.crit)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 2h. 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
Fig. 3 Test-equipment
0.2 360
,J
-Detail of notch region
see aLso
Fig. 2e Test piece P
Conditions for maximum thickness t = 65 mm
When t > 65 mm the total height of the test piece should be t+2/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 O.O7t. Fig. 2a Test piece T and the way of loading (proposed for acceptance testing)
Machined flush with plate
\(both sides)
See detail
\
Center of weldL(1t Il H Il
!LI U Wittl±L
lo
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 a view to the reproducibility of the test results (fi-gure 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 ± I0%. (See appendix I.c). In figure 4 a device is shown which allows quick and safe manipulating with the drop-ham mer.
Fig. 4 Magnetic safety clutch (when current is switched on.
pin in magnet goes down)
Before subjecting a test piece to drop-weight loading the notch has to be prepared for the measurement of the residual C.O.D.
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.
Very reliabl, results (Adjusting of pin of
extensometer easy.)
Fig. 5 Possible methods of measuring COD.
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 is the residual C.O.D.
Table I. Dropheight sequence
Figure Sa A A
Cheap specimen
(Pin of extensom,tpr
adjusted from
origi-nal circutar section to indicated section. This is difficuLt to do without impairing the accuracy.) Hl
H2 H3 H4 H5 H
H7 H8 H9 mmm 250 300 350 400 450 500 550 600 700 etc.The second blow is given from a height of 300 m (see table I), 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 C.O.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 C.O.D.' s are plotted with the temperature ordinates as a base.
Blow nr. 8 550 500 5 450 4 4 400 3 350 2 300 i 250 o c.o.D. Fracture> 0.18mm - ist; 3rd blow not
mea-Fig. 7 Maximum C.O.D. 's 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.
lt 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 maybedecisiveforthequality of the weidment and may be determined in a similar way.
4 Determination of critical C.00D. 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 = M/ W is:
P = 1 la,, t a: kg/mm2 (yield-stress) t: mm (plate-thickness) P: kg (load)
In figure 8 not only the C.O.D. at incipient yielding but all 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 the static 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 still 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 a 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). Therefore the much more accurately to determine total C.O.D. at incipient yielding was chosen; it is about equal to the residual "plastic hinge" C.O.D. (see figure 8) for the test conditions chosen.
- -
-0.09; .o Fracture 1racture F!thr!Jao5am.0:1mm
::
Fracture 0.031 0.033 0.028 mm 0.027 10025 10.016 0.022 . I Atte and 0.013 0.012 0.007 0.007 lP.008 I C.O.0 J ours 0° +100 Test temperature (°C)-. Semi-killed steel 37Isotherm Robertson arrest temp. ±20°C
Charpy 3.5 kgm/cm° (20 ft.lb) ± 3 oc 5.2 kgm/cm° (30 ft.lb) ± 7°C 7 kgm/cm° (40 ft.lb) + 9°C
50% crystalline +17°c
Fig. 6 Example of testing diagram for given mild steel 600
700
600
12 4500-4000 3500 1.sG.y 3000 2500 2200 E 2000 15 13 12 11 10 g B 7 No fracture at C.O.O_0.55n-,m 16 - -
--- °C
No fracture at CODs 1.5mm cul _15°C 14//
,4-j 4. Ca leu tao en 0 2rt' saw cut VISUa ->Ptastic hinge minimum .2 specified proof StressCOD. at incipient yield to be
used as criticaL value for residual CO.D in dropweight test.
Residual COD. after unloading
from load at which plastic hinge
is formed.
Semi-killed steel 37
Isotherm Robertson arrest temp. + 20 oc
Charpy 3.5 kgm/cm (20 ft.lb) + 3°
5.2 kgm/cm° (30 ft.lb) + 7° 7 kgm(cmi (40 ft.ib) + °c
50% Crystalline +17°C
Fig. 8 COD. values obtained with static bend-tets at 5°C and 50°C for given mild steel 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.
5 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-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 weidmetal 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 1f 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
u 1500 s 1000 3 500 E z
0
0 J Visual/ Calculated Incipient yielding
0.2
COD.
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 COD. 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 T-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 I.- versus T-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 oft/ic 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 COD. 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 [I] 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:
14
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.1.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. lt 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
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 weidment (figure 10). A third reason is that the test-equipment turns out to be simple and in expensive.
Fig. lo 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 Ib). 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 os/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
de-16
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. lt 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 CO,-welds are collected in Appendix Il. 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.J.W. Commissions II, IX, X and Xli 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.
References
AuDIGE, A., Progress reports of Working Group ,.Brittle
Fracture in Service". 11W-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, II W-doc. 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-doc. IX-573-68/X-567-68.
IiOEDA, K., Y. AKITA and H. KIHARA, The Deep Notch Test
and Brittle Fracture Initiation. 11W-doc. X-404-67.
NIBBERING, J. J. W., Plastic deformations at notches in welds
of mild steel plates. S.S.L. rep. 129, (11W-doc. 2912-107
1968).
NIBBERING, J. J. W., J. VAN LINT and R. T. VAN LEEUWEN,
Brittle fracture of full-scale structures damaged by fatigue. Neth. Ship Researchcentr. Report no. 85S, 1966. 11W-doc. X-374-66 and 1.S.P. Nov. 1966.
VAN DEN BLINK, W. P., A crack extension test for weld metal. II W-doc. IX-527-67/X-453-67.
MARQUET, F., Side bend test procedure. Steel times, Nov. 19,
18
APPENDIX I
a. Influence of drop-height of first blow on test results b. Influence of step-magnitude on test results
.-Fracture 7 :1 Blow C.O.0.(mm) 500 I , nr 0.049 0.054 450 S 0.038 0.045 400 l---4 0.027 0.033 550 O 350 E 300 E .0 O' o. o 200 100 O C.o.D. 3 0.017 0.025 2 0.009 0.016 0.008 Tent temperature O °C Average of 2*2 specimens C.O.D
c. Influence of variations in drop-weight on test results
Fracture 1000 : Blow COD. (mm) nr. Fracture 500 . 1 0.010 Blow C.O.D.(mm) 460 r 0.040 400 4 0.031 0.027 0.012 250
°
1 0.007 Test temperature: -5°C Average of 22 specimens Conclusions:Influence is moderate, (about l0% of critical C.O.D. for a
difference of 50 mm in
ini-tial height)
Conclusions:
Results conform if plotted
on the basis of energy
(weight x height). Difference
in fracture C.O.D. is not
alarming (10%)
I
f >Fracture F /Blow C.0.D.(mm) nr. 350 -. 300 250 200 Test temperature O °C Average of 2o2specimens COD. 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 e, 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 1 length of b. 50,000 0.4 ±10 25 9 0.240 fatigue-cracks c. 96,000 0.5 ±20 30 5 0.360 0.5 mm d. 71,000 2.5 0 30 1 0 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 3.5 +20 15 6 0.037 1 mm i. 101,500 1.8 ±30 15 6 0.030 I j. 110,000 ±2.5 +40 15 6
080)
3-6 0.041 0.054 5 0.045 2-4 0.020 0.033 3 0.025 1-2 0.007 0.016 1 0.008 5 0.046 4 0.033 0.026 3 0.015 900 -800 700 600 600 550 500 450 400 350 300 250 200 ob Submerged arc weld: L-test
Symm. double V, 90°, two passes, resp. 800 and 1000 Amp.
Weld speed 33 cm/mm, wire 5 mm, cr = 35,8 kg/mm2, = 48,5 kg/mm2
Temp. C.O.D. mm
0°C 0.19 critical temp. - 15°C (estim.)
- 5°C
0.09 (Charpy 3.5 kgm/cm2:10°C
0.065c Submerged arc weld:
T-test; with lack of penetration (see figure 9) Temp. C.O.D. mm
APPENDIX Il
L-test; symm X; 60°; hor; root pass: basic electr. Composite electrode (basic core) 2.4 mm
450 Amp; 45 cm/mm; 8 passes Temp. C.O.D. mm
40°C
0.04130°C
0.05520°C
0.02015°C
0.2410°C
0.63 critical temp.: J - 18°CI°i
(Charpy 3.5 kgm/cm2 20°C )Summary of tests on welded plates (Dropweight 22 kg)
a Plate material: (Semikilled, as rolled)
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 thickness 25 mm; yield point 25 kg/mm2
tensile strength 45 kg/mm2 - 10°C 0.11 critical temp. J - 10°C J
- 10°C 0.165 (Charpy3.Skgm/cm2 :J+25°C!)) critical C.O.D. 0.055 mm
10°C 0.050
Results: (- 10°C 0.045)
-
fracture at copper inclusionTemp. C.O.D. mm 0°C 0.400 next to notch)
20°C
0.01 critical temp. J- 10 °C 0°C 0.140 - 10°C 0.045 (Charpy 3.5 kgm/cm2:-
7°CI) 0°C 0.190-
10°C 0.070 50% cryst. +7°C0°C 0.280 y. d. Veen + 8°C)
e Automatic CO2 weld:
30°C
0.160 critical temp.30°C J
20°C
0.150 (Charpy3.Skgm/cm2:f Electrogas weld (enclosed CO2): T-test
Composite elestrode 2.4 mm; 450 Amp; 5 cm/min Temp. C.O.D. mm
10°C
0 critical temp. : + 5°C J (estim.)0°C 0090 (Charpy3.5 kgm/cm2: - 20°C J) 0°C 0.035
PUBLICATIONS OF THE NETHERLANDS SHIP RESEARCH CENTRE TNO
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I S The determination of the natural frequencies of ship vibrations (Dutch). H. E. Jaeger, 1950.
3 S Practical possibilities ofconstructional 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
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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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9 S Stability experiments on models of Dutch and French standard-ized lifeboats. H. E. Jaeger, J. W. Bonebakker and J. Pereboom, in collaboration with A. Audigé, 1952.
10 5 On collecting ship service performance data and their analysis. J. W. Bonebakker, 1953.
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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.
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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
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(Dutch). J. C. G. van Wijk, 1957.
19 M Crankcase explosions (Dutch). J. H. Minkhorst, 1957.
20 5 An analysis of the application of aluminium alloys in ships' structures. Suggestions about the riveting between steel and
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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
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25 S Experimental determination of damping, added mass and added mass moment of inertia of a shipmodel. J. Gerritsma, 1957. 26 M Noise measurements and noise reduction in ships. G. J. van Os
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27 S Initial metacentric height of small seagoing ships and the in-accuracy and unreliability of calculated curves of righting levers. J. W. Bonebakker, 1957.
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 5 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 momeiis for three mod-els of different fullness in regular waves. J. Ch. de Does, 1960.
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41 S Stresses in flat propeller blade models determined by the moiré-method. F. K. Ligtenberg, 1962.
42 S Application of modern digital computers in naval-architecture. H. J. Zunderdorp, 1962.
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45 S Some acoustical properties of ships with respect to noise control Part II. J. H. Janssen, 1962.
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A. M. van Londen, 1962.
47 C Results of an inquiry into the condition of ships' hulls in relatior to fouling and corrosion. H. C. Ekama, A. M. van Londen ant P. de Wolf, 1962.
48 C Investigations into the use of the wheel-abrator for removin
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49 S Distribution of damping and added mass along the length of shipmodel. J. Gerritsma and W. Beukelman, 1963.
50 S The influence of a bulbous bow on the motions and the propul. sion in longitudinal waves. J. Gerritsma and W. Beukelman, 1963 51 M Stress measurements on a propeller blade of a 42,000 ton tanken
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52 C Comparative investigations on the surface preparation of ship building steel by usingwheel-abrators and the application ofshop coats. H. C. Ekama, A. M. van Londen and J. Remmelts, 1963
53 S The braking of large vessels. H. E. Jaeger, 1963.
54 C A study of ship bottom paints in particular pertaining to tin
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 Curaçao
Dutch Lesser Antilles, P. de Wolf and M. Meuter-Schriel, 1963 57 M Determination of the dynamic properties and propeller excitec
vibrations of a special ship stern arrangement. R. Wereldsma
1964.
58 S Numerical calculation of vertical hull vibrations of ships b
discretizing the vibration system. J. de Vries, 1964.
59 M Controllable pitch propellers, their suitability and economy foi large sea-going ships propelled by conventional, directly couplec engines. C. Kapsenberg, 1964.
60 S Natural frequencies of free vertical ship vibrations. C. B. Vreug
denhil, 1964.
61 5 The distribution of the hydrodynamic forces on a heaving an pitching shipmodel in still water. J. Gerritsma and W. Beukelman
1964.
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 diese
engines using heavy fuels. R. W. Stuart Michell and V. A. Ogale
1965.
64 C Barnacle fouling on aged anti-fouling paints; a survey of pertinen literature and some recent observations. P. de Wolf, 1964.
65 S The lateral damping and added mass of a horizontally oscillatini shipmodel. G. van Leeuwen, 1964.
66 S Investigations into the strenght of ships' derricks. Part I. F. X. P
Soejadi, 1965.
67 S Heat-transfer in cargotanks of a 50,000 DWT tanker. D. J. vai der Heeden and L. L. Mulder, 1965.
68 M Guide to the application of Method for calculation of cylinde liner temperatures in diesel engines. H. W. van Tijen, 1965 69 M Stress measurements on a propeller model for a 42,000 DW
tanker. R. Wereldsma, 1965.
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 bo W. P. A. van Lammeren and J. J. Muntjewerf, 1965.
