Fracture mechanics and fracture control for ships

ARCHEF

LABORATORIUM VOOR

SCH EEPSCONSTRUCT1ES

TECHNISCHE HOGESCHOOL

-

DELFT

RAPPORT Nr.

SSL 167

BETREFFENDE:

Fracture 'mechanics and fracture control for ships.

FRACTURE MECHANICS AND FRACTURE CONTROL FOR SHIPS

by

Prof.ir. J.J.W. Nibbering

Ship Structures Laboratory,

Deift University of Technology, Mekeiweg 2, DELFT,

The Netherlands.

Report no..

CONTENTS: Page

General 15-1

1.1. Defects

1.2. Residual stresses

1.3. Fatigue 15-3

1.4. fligh heat input welding

Acceptance testing and fracture mechanics

for ships 15-6

11.1. Crack-arresting

11.2. Welding material 15-9

11.3. H.A.Z. t!

11.4. The relative importance of metallurgical factors and crack-length

II,. 5. Critical C.O.D. 's and impact testing 15-10

11.6. C.O.D.-measurements 15-14

11.7. Final observations 15-15

15-1

FRACTURE MECHANICS AND FRACTURE CONTROL FOR SHIPS.

By _{Prof.ir. J.J.W. Nibbering.}

Cracks in ships are either

fatigue-cracks or brittle cracks.

Fatigue-cracks are most _{common but the danger involved}

is ma1l.

Brittle cracks are _{very scarce nowadays, but when}

they occur, the ship ob-viously is in danger.

Both types of cracks _{generally start at defects due to welding}

or flame

cut-ting especially when _{these defects are situated}

at geometrical stress

con-centrations.

1.1. Defects.

In ships with their _{enormous amounts of welds (between}

10 and 1000 km in

length), weld defects _{can never be completely avoided,}

despite intense

non-destructive control.

The most important defects

are under-cuts, lack of fusiàn, _{slag-inclusions,}

incomplete penetration and _{cracks in welds and heat-affected Zones.}

In princip1e one _{can achieve that not any defect}

will develop into a large

brittle fracture. For this,it _{is necessary that}

proper weld and parent

mate-riais are chosen and welding _{methods are avoided, which}

excessively destroy

the originally sound _{parent material in a relatively}

ectended zone. Even the

mere use of parent material of

such high quality that any eventual brittle

crack will be arrested, is _{mostly completely satisfactory.}

Then the quality

of the welds and the _{H.A.Zones would be of}

secondary importance from a

safety-point of view, and this could lead to substantial

reductions in cost

of welding, production _{and quality control.}

Unfortunately when high heat

input-welding methods like _{one pass submerged-arc, electrogas}

or

electro-slag welding is applied,

eventual cracks starting in the _{welds or}

heát-af-fected zones, _{are not always leaving the}

welded region under the _influence

of the residual welding

stresses, as is the case with multipass-welding. (See 1.2). 1.2. _{Residual stresses.}

The stress field set _{up with high heat input welding}

has a much smaller

gradient than the _{one created by low heat input}

welding, (fig. 1). Due to

that the shear _{stresses in planes parallel}

to the weld are also _smaller,

and so are the _{corresponding normal stresèes,}

of which the direction _differs

from the weld line.

As fracture will always _{develop in a direction}

perpendicular to the main

ten-sile stress, it will _{deviate more from the weld}

line, the higher the shear

stresses be.

In figures 2a and 2b _{the stress situation for}

a small element situated in the H.A.Z. is indicated;

an estimate of the angle a between the load

stress-es (nominal strstress-essstress-es) and the _{main tensile stresses is}

.niade in fig. 2c and

2d. It should be reminded _{that the shear stresses}

'u Gioad FIG.

LOW HEAT-INPUT

Gres

Tres

F1a2c

LARGEST MAIN STRESS

.fDEVIATES

o WITH

LOADING - STRESSES 15-2 FIG.1 1Tres. Gioad Gjoa

CC-.-arctg

2tres.

-

GioadG.es

FIG.2b

HIGH HEAT-INPUT

Gres.

Tres.

ires.

Load

Les.

LARGEST MAIN STRESS

NEARLY IN LINE WITH

LOADING - STRESSES (o SMALL)

5-3

the residual tensile stresses (ares) also vary along the weld line. This is valid for hand-welding with many stops and starts, but far less true for

automatic and semi-automatic welding, especially with welds _{made in one pass.}

Residua], stresses have often been blamed for causing brittle fracture.

This was right for ships constructed during _{and shortly after World War II,}

when low-stress brittle fractures were very common and the steel was often

not good enough to arrest them. But nowadays residual _{stresses are rather}

beneficial. The ships' steel is so much better than in the old days, that

eventual cracks started in a weld or H.A.Z. _{are mostly arrested in the}

par-ent plate before attaining a critical length. (This is not the critical

length under static conditions, which is in the order of meters, but the

critical length belonging to the high-speed state of loading -.curring when

a brittle crack propagates).

1.3. Fatigue.

The residual stresses have also a great influence on the propagation of

fatigue-cracks: partly beneficial, partly detrimental.

The beneficial effect is again that cracks are forced to leave the welded

region. The detrimental effect may be that cracks initiate _{and propagate}

easier when residual tensile stresses _{are perpendicular to the crack front.}

This is generally the case with cracks developing _{perpendicular to a butt}

weld. In other cases, for instance when cracks develop (more _{or less)}

par-aile]. to a butt weld, the propagation may be slowd down. (Of course the heterogeneity of the H.A.Z. also plays a role).

All these

influences

makè the use of a simple crack propagation _{formula like}

da

m

very unreliable. The value of m can vary between 1,5 and 6 /1/. _For

ship-building, problems like the random character of the loading, _combinations

of in-plane and normal-to-the-plane loading and corrosion add to the

corn-plicacy of the problem.

On the other hand it is fortunate that a. high accuracy in estimating fatigue

crack growth is not important for ships. One can hardly speak of a critical

crack length, because the notch toughness of shipbuilding steel is good

enough to be able to bear the presence of really large cracks. Leakage there-fore is a more important restricting factor. But as most cracks develop in

internal stiffening members, leakage often can be neglected.

It has been emphasized that due to the, inevitable presence of defects, cracks

soon develop under the action of the cyclic loading and consequently the

re-sistance to _{ropagation of cracks is mainly determining the fatigue life of}

ships' structures. Nowadays this is even more true, because of the very

large dimensions tankers, bulk carriers and containerships have obtained.

A crack of 10 cm depth in a 50 cm high stiffener of a small ship reduces the

cross section as iiuch as a cm crack in a 2 m high stiffener in a large

ship. Yet the time for development of the latter is about two times as long

as for the former.

Fatigue-cracks are not only relatively harmless _{as such, they can even be}

beneficient. For, whenever brittle cracks develop in ships, they always

28

24

20

'I,

3Srnrn(500C)

1h. dIdsnc. from V-notch to fustoi line and 1h. peak-temperatur, at that pleca

., during welding ura indicated

at the cwa. emm(m20c) tat. m.tetlal (unaffected) -

p -.

...-... -80 -50 -40 -30 -20 -10 0 +10 TEST TEMPERATURE (.°C) --70 15-4

this of course is often the _{case at the tip of weld- or H.A.Z.-defects.}

(See section 2). A little extension of these defects will replace the tip from the material in bad condition to that of normal state, and initiation

of a brittle crack becomes far _{more difficult.}

1.4. High heat input welding.

There are however some points which need special attention. Firstly again the case of high heat input welds. It is very well possible, and

demonstrat-ed in /1/, that fatigue-cracks _{propagate completely in the. weld or H.A.Z.}

In the case considered a 3 mm wide (or narrow!) _{coarse grained zone existed}

of very poor notch toughness along _{an E.G.-weld in a 34 nun plate.}

Notwith-standing the narrowness of this _{zone, high stress - low cycle fatigue cracks}

indeed propagated in that zone over quite a distance (up to 120 mm). The

large danger involved can be appreciated when it is known _{that the}

fatigue-crack several times "jumped" forth in a brittle way over a small distance (5-10 mm) before it finally developed into a complete brittle fracture which also ran ail along the weld line! The test-temperature was -20°C, the nomin-al fracture stress 24 kg/mm2. The steel was extremely good (E-qunomin-ality, as

is obvious from the Charpy V notch _{energy which was 16 kgm/cm2 at -60 C!}

The reduction in quality in the H.A.Z. due to the high heat input of the E.G.-welding is shown in fig. 3.

FiG. 3 _{CHARPY-V VAWES OF THE (AL St 52 Nb. (THICKNESS: 34mm)} AND THE LG-WELD. (5]

Section weakened

_{by s/t.}

15-5

In terms of "safe"

temperature the steel has been spoiled _{nearly 100°C.} Similar dangerous

situations can also occur with more normal _welding meth-ods, - especially when the number

of passes is not very high - _{, or when} one side butt welding has _{been applied. Fig.}

L _{shows some examples.}

FIG.4

FIG.5

It will be clear, _{that cracks developing}

at the points of _incomplete penetra-tion will not easily

leave that vertical plane, as long as the cross-section remains reduced by that lack of

penetration. A similar situation may occur at undercuts and long lacks of fusion, (fig. _5).

It is evident that

transverse butt welds in ship decks, sidés _{and bottom are} more critical in this _{respect than longitudinal}

butt welds. But generally only really

_¿

_{defects are dangerous.}

Most of the standards _{and specifica.i.} tions as to defects are too pessimistic.

On the other hand it is well recog-nized that these

specifications have mainly the purpose of _{guaranteeing a} satisfactory level of workmanship,

thus have a function in quality _control which of course _{can never bè dispensed}

wïth.

15-6

II. Acceptance testing and fracture mechanics for ships.

The currently used ships' steels and welding materials have such good tough-ness that the application of linear elastic fracture mechanics and the use

of K1 or G1 as. minimum required toughness values for static loading have

no sense. This is partly due to the severe requirements set for these

mate-rials, especially because not static but 4pact tests are prescribed. The second cause is that the temperatures ships are subjected to are not

too low, connected to the fact that the lowest sea temperature is about - C.

Finally for ship steels metallurgical influences, due to welding and flame-cutting overshadow largely the influence of geometrical factors like

crack-length. (See II.').

11.1. Crack-arresting.

Nevertheless when it is required that ships steel in deck and bottom should

be able .to arrest any brittle crack, fracture mechanics can be of some use.

But the most realistic way for estimating the involved dynamic toughness of the material (KID ) is by carrying out a crack-arrest test of the Robertson

type, and evaluat it for the maximum length of crack to be arrested. (That

length may be appreciable because a crack started in the side plating, may

have developed, many meters before it finally meets the sheer strake in which

it should be arrested). However, when looking to the usual results obtained from isotherm Robertson tests, it always shows a tremendous increase in dynamic toughness over a very short temperature region; for instance 5-fold KID over 10°C, (fig. 6).

As K1 is about proportional to the square root of crack length, a 5-fold increase inK1 would for a certain temperature mean a 25-fold increase in critical crack length. In other words: the sensitivity of the material to

temperature is much larger than to crack length.

In quantitative terms: when a material has an isothermal Robertson crack arrest temperature of -10°C, - which means that cracks up to about 30 cm

can be arrested - , then. will it be able to arrest cracks of 5 m - 10 m

length at 0°C. In such a situation it is.of course wisest to require a Rbertson crack arrest temperature of say -15 C and take it that it may be

5 C too safe. Most steels of D and E quality are certainly mOre "safe".

It is unfortunate that in the specifications of steels reliable crack arrest tests are not required. The desired properties are more or less obtained by requiring certain Charpy-energies at certain temperatures. The correlation

between these transition temperatures and the Robertson crack arrest

temper-ature is not only poor, (the latter is on the average some L0 C higher than

the 3,5 kgm/cin2 Charpy temperature), but the scatter is also large. (For individua], steels the difference between 3,5 kgm/cm2 Charpy and Robertson

crack arrest temperature can be 0°C to 80°C. (See fig. 7, obtained from /2/

by Di's. H.C. van Elst)). This does not mean that the Charpy-test has not any use in this respect. Verbraak and van Elst have already long ago propag-aûdized to use it for quality control for steels with specific composition, made according to a fixed procedure. For such a well defined case, the

rela-tion between the Robertson crack-arrest temperature and for instance the

Charpy-energy at that temperature shows little scatter and can be used for

quality control. n, the same way the Charpy test can be used for quality control of welding material.

L

L

C

15-7 16 20 12

-100

ARREST

)

THROUGH ÖRACK CRITICAL

X N.C.R.E..

O STEELMAKERS

CONSTANT TEMP TESTS<9AKERS

6 FT. WIDE PLATE TESTS.

.

IO

-80

-60

_-40.

Temperature (°C)

o.

z'

-20

FIG.6 RESULTS 0F

_{CRACK-ARREST TESTS.}

(FROM ADMIRALITY REPORT [4]

).

X.

J

15-8

1o.

<o

o -o

O 2

4b

3,5kgrn/crn2 (=2Oft.Lbs.) Charpy-.v temperature (°C)

FIG.7 RELATION BETWEEN

_{20 ft.Lbs. CHARPY-TRANSITION}

TEMPERATURE AND CRACK-ARREST

_TEMPERATURE.

(van ELST. [2] ).

r

40

20

o

o

o.

o

o

o0

-4O

_o

0/

6O

15-9

11.2. _{Welding material.}

In the I.I.W. the W.G. 2912 has the task to develop reliable quality

con-trol and acceptance tests plus criteria for welds. During the chairmanship of van den Blink, the group has come to the opinion that for weld material

crack arresting is of no importance, _{because as mentioned in section I,}

cracks always tend to leave the _{welded region.}

The opinion was based as well on overwhelming practical experience _{as on}

results of experiments from Kihara and _{Ikeda, (see /3/).}

The conclusion was that for welds only the resistance to crack initiation was of importance.

If so, it would be realistic to stipulate less severe requirements _{for welds}

than for plate material, because _{the latter had to be resistant to crack}

propagation which is a severe dynamic phenomenon. (Of course initiation _of

cracks can also occur as a consequence of dynamic loading, but Seldom _as

severe as occurs during brittle crack propagation, of which the speed is

some 2 km/sec.).

11.3. H.A.Z.

Another point is, that often the _{H.A.Z.-properties need not be controlled;}

yet they may be worse than those of the weld metal. The inspection boards

meet this by having in their (Charpy) _{requirements for plate material built} in a safety margin, which allows

for a certain deterioration of the plate materia.], in the H.A.Z.

One might regard upon it in this _{way: when the plate material is resistant}

to brittle crack propagation, the H.A.Z. will _{at least be resistant to the}

much milder crack initiation. This method may be right for normal,multi-pass welded steels; it is wrong for normalised or quenched and tempered fine

grain steels, welded with high heat input.

This has already been demonstrated in figure 3. A logical approach _{would be}

taking many Charpy-specimens from all over the H.A.Z. and stipulating _a

low-est acceptable value. This indeed occurs more and more, but unfortunately

not always realistically. For instance 6,2 kgm/cm2 _{at -30°C as required by}

some inspection boards for ships, is much too severe.

In the Delft Ship Structures Laboratory full scale éxperiments with 34 nun

thick E.G.-welded plates of Nb-normalised steel have proved that aminimum value of i kgm/cm2 Charpy-energy at the lowest service temperature would be

sufficient. (See /5/ and below section _11.4).

11.4. _{The relative 3. ortance of metallurjca].}

factors añd crack-len: h.

In the discussion in 11.1 on crack-arresting, it has been discussed that _the

influence of temperature on the toughness. of a material subjected to high

speed loading is so large, that crack length is only of second rate

impor-tance.

When discussing crack-initiation, _{the situation is similar in}

very brittle conditions. These conditions _{may exist in H.A.Zones. As an example is again}

chosen the H.A.Z. of _{an E.G.-welded Nb-nortnajised thick plate.}

In figure 8 the specimen is shown, and in figure 9 some results are given.

From the numbers near the points it can be sen that crack length had not

r-e

e

15-10

WEI.D PROC 0AiE FATIOI SPECI4ENS: TISCICNESS U.v,i.

WELDS j)® ® ®® AND (iJ ARE E1.ECTROGAS WELDS W 1I.. PtATES Ai ELECThDSI.AR ii ie PLAiES WELD(J EUBMEPOED- ARC WELD

THE NDTCMEE ., F ARC F2 WERE MACE AFTER WELOUAi. AI.

AND REFØRE.WELDIPW ® IN,W-NOTCN).

THE OTHER NOTCHES WERE MADE AFTER WELDiNG

FIG.8 [I

i

8

any inflüence on the results. But the distance from the crack-tip to the

fusion line was of utmost importance: a difference in position of i nun had

a greater influence than a difference in crack length from 26mm to 100 min. In fig. 9 both influences do not appear as clear as possible. A presentation of C.O.D. (see 11.6) as a function of temperature, as in fig. 10, is much more discriminating than fractuSe load versus temperature. Following a ver-tical line in the region of -25 C crack lengths increasé from 26 to 36, 53,

76, 1011 and 79 mm's. This is quite opposite from what fracture mechanics would predict. But another-look shows that the lowest points generally

be-long to cracks situated on, or close to the fusion line (sis) and the higher

points to cracks further away. Together with fig. 3 -this demonstrates the

large differences in material quality in a very narrow region. The figure also shows the enormous influence of temperature on the CIO.D. to fracture:

100C rise in temperature raises the C.O.D. from 0,015 to some 0,5 nm,

1/3 (point

11.5. Critica]. C.0.D.'s and impact testing.

In such cases it has of course little sense to discuss thoroughly which

value of C.O.D. should be required as an absolute minimum. Whether it is

0,1 mm. or 0,2 smi does not make rea]. difference in structural quality. On the other hand, for autothatic welds much smaller increases of C.O.D.

with temperature have been observed. (Fig. 12 /6/). Therefore it has been

proposed to require always at least 0,3 min C.O.D.0 in static testing.

This will only

in exceptional cases be too much on the safe side, but for most cases it will be as good as a slightly different value; it has however

- the advantage of being easily measurable.

It should be underlined that static C.O.D.-testing is not always, or even seldom, sufficiently realistic. Many structures may be subjected to regular

42 41 40 39

3e

37 36 34 33

i

SA.-w.Id/3_{(T.niII. tut aftur fats-LaudIng)}

104 32 31

-£ 5/i ¿76 79

15-II

1/3

4

(T,n(i tut aft

HIGH STRESS I FRACTURE sie £ St52.N E.-w.tdsd. (Thlcknsu:3immJ

0+ XD1

COMPLETE FRACTURE NUMBERS AT POINTS: f DISTANCE TO FUSION-LINE

AND TOTAL CRACK-LENGTH NOTCHED .WELOED

i a i t

-27 -26 -25 -24 -23 -22 -21

-20 -19 -18

-17

-16 -15

-14

Test temperature(°C)

FIG. 9

_{NET FRACTURE -STRESS AS FUNCTION OF TEMPERATURE. C5]}

-13

-12

-11

rH

E .26 C

o

u

4,

u

C a-4,

o

UI UI 4, L.

30

29

28 27 26 25 24

23

22'i

1/3 - S5(N.W)

'26j'/3

5ß'IlIrs..NoT 26 26 41/3 626 25 26 FATIGUE-LOADED (SEE Fia 20) ¿3/3 ¿26 _40/0

+10

26 LOW STRESS FRACTURE (SEE FlG.21) 0.5(.iW 26 1/3 27 1/3 4 PLATE 26 I SPECIMEN 1 (n-34000)

io - -

2 (n-O)

21

-4,

z

₂₀

i t i i i i a i i « z' (n.1500) 3/3k FICT_1*

_'r

_{12 (n.660)} L

"

12' (n.5000) 6/6 65 PLATE 26 3 53 1/3 26

80 10 9 8 7 6 5 4 3

12

E 1

o

-

Ss&-wstd/3

(TENSILE TEST AFTER FATIGUE-WADING)

101

(SEE APPI -FIG 2)

/

el

/

2.5

/

75 (SEE Fl21) 26 (Nat fatigus-toadid) °h° 1/3

:5®

O.5(N+W) , (SEE Ff320)

--cl

'/3k

t

1/3 35 ¿ 2.5(N.W) 3,3 . 442 3

+26

¿0 4

-1o/0

SI 28 1/3 79 35tN.w 26 bio 26 3 _1/3 53 ₅₀ Test temperature (°C) 15-12 It11

t

9/6 .5

14

f/s 69 PLTE

/

I

L,

PLATE 26

I

¡

I

TENSILE TEST AFTER FATIGUE-LOADING)

St82 +N E.G-imld.d Thlckn.u: 31mm. SPECIMEN i (n31OØ0) O - 2 (n.0) MO + 2 In-1500) FRACTURE _{- fl (n.660)} O 12'(n.5000) 29 (n.10000) * 1.5 mm 9/6 94 COMPLETE FRACTURE r PARTIAL FRACTURE (SPECIMEN. 1) NUMBERS AT POINTS DISTANCE TO FUSION-LINE AND TOTAL CRACK-LENGTH.

(N.W): NOTCHED .WELDED

O t I I I I I I I I

I

I I I I I

-28 -27 -26 -25 -24 -23 -22 -21

-20 -19 -18 -17 -16 -15 -14 -13 -12 -11 -10

FIG. 10 DEFORMATIONS AT THE NOTCH 'S _{TIPS AT THE MOMENT THE FRACTURE} OCCURED GIVEN AS FUNCTION OF TEMPERATURE.[.5J

-H

-D

70 40 30 20 18 16 14 12

1,5 1,4 1,3 1,2 1,1 14,0 01,9 o,e 0,7

jO,6

d

0,5

Ò

i:

0,2 0,1 O

70-60

15-33

Test specimens according to OflA.commIttee.

pLate

thickness

£

46mm

_(Es.)

- 34mm

(E.G.)

D ---

- 22mm(A)(E.Q)

22 mm(B) (E.G.)

-50 -40 -30 -20

Test temperature ()

FIG.12

_{C.O.D. AT FRACTURE}

_{VERSUS TEMPERATURE.}

(ELECTRO -SLAG

_{AND ELECTRO-GAS WELDS).}

I

-I

r'

I j

15-14

In view of this it is fortunate that most inspection boards define their specifications in terms of energy obtained from Charpy-impact tests.

What difference this makes for weld metal is evident from table I, column 3

and 5.

TABLE I. /6/

The main shortcomings of the Charpy-test are: reduced plate thickness;

small, rather blunt, notch;

recorded. energy is sum of initiation and propagation energy;

instead of energy, the deformation at the notch root (C.0.D)

should be measured.

These objections have been avoided in the so-called Niblink-test, which actually is studied in the I.I.W. in W.G. 2912. For information, some

re-suits are also given in table I.

11.6. C.0.D.-measurements.

The attractiveness of the C.O.D.-concept is that its physical meaning is quite clear: it is the deformation (extension) of the material at the tip

of a notch or crack. Wells /7/ has introduced the concept.

For lower strength steels of moderate thicknesses it is much more realistic to specify required fracture toughnesses in terms of critical C.0.D.'s than

in terms of K or G These steels thank their structura], usefulness

Ic Ic.

mainly to their large plastic deformability, so that a ductility-requirement

is a logical consequence.

As discussed before critical C.0.D. 's for these steels should be in the

ord-er of magnitude of 0,2 to 0,11 mm.

Even for cracks as large as 100 mm in length, these values can only occur

at nominal stresses equal to or larger than yield point.

This can be illustrated as follows.

For a crack with a plastic zone 2r _{, the virtual crack length is about}

2(a+r).

y

r

Type of weld Thickness in mm. - Charpy

3.5 kgm/crn2 Niblink test dynamic 0.06 mm. C.0.D.A. test static 0.3 nun. ES 46 +10°C +20°C +22°C Sub.Arc 116 - - 11°C -46°C EG 34 +13°C +18°C -19°C Sub.Arc. 34 - +13°C -10°C EG (A). 22 - 8°C - 3°C -62°C Sub.Arc. (A) 22 -180C -13°C -23°C EG (8) 22 +73°C _{+27°C -24°C} Sub.Arc. (B) 22 + 5°C - -63°C

15-15

From E. (a2 - x2 _{the C.0.D. at crack tip can be calculated by}

sub-stituting a + r for a and a for x +

COD

r,2 +2a.r; when

_{C.O.D. = 0,4 and a}

30 kg/mm2 + r

- 35 n.

When linear elastic fracture mechanics would be valid for this _situation,

the nomina], stress should be _{equal to the one calculated from}

aa2 20 2

-For a steel with a yield point _{of 35 kg/mm2 this leads to}

f2r.a2

I

a =

y --/ 2.35.352 ₂

a

_-v

50

:41kg/nun.

This is indeed larger than yield point, _{and linear elastic fracture} mechan-ics do certainly not apply.

For a C.0.D. of 0,3 mm, r 22 mm and a _{33 kg/mm2. This is also too}

y caic.

close to _{for valid calculation results. It}

may be questioned whether COD

is independent, of crack length.

From

COD _{r2 + 2a.r}

-it can be seen that when a _{a,, the larger the crack, the larger}

the C.O.D.

at a certain size of plastic zone. Now, as is known /7/, the size of _the

plastic zone (in relation to pláte _{thickness) is a measure of the}

stress

state in that plastic zone, (r _{t + fully plane stress). Consequently}

the larger a. crack the larger the C. O D _{needed for plane stress.}

Nevertheless, in view of the _{many unknown factors involved,\ and for cracks}

or defects between 10 and 100 mm length, _{one single C.O.D.-value will} suf-fice. It has.the advantage that _{for very thick plates, Wells' requirement} of r _{t does not lead to unrealistically -high toughness values. A C.O.D.} of say 0,3 mm indeed represents

a quite satisfactory deformability. The fact

that to it corresponds a situation of plane strain in very thick plates is

not objectionable at all.

11.7. _{Final observations-.}

-Figure 10 is a fine illustration _{of the usefulness of the C.O.D.-concept.}

-There is a clear separation between _{the "bad" and "good" results, which} con-trasts sharply with fig. -9. The figure however demonstrates also how

impox'-taüt welding parameters may be, and that these may completely overshadow _the

influence of other

parameters like crack length.. This has already been dis-cussed before as fax' as the position of the cracks relative to the _fusion line concerns. -But in this final paragraph the attention is drawn to such an

apparently secondary factor like _{sequence of welding. The point quite to the}

- right: O (N + W)

- represents a partial fracture, which started

at

- 28; n =10.000

-15-16

a notch1 which was made in the H.A.Z. of a transverse E.G.-weld prior to the subsequent longitudinal submerged arc welding. The distance between the

notch and the S.A.-weld was so small that, the material at the notch tip

du-ring the S.A.-weldig was strained plastically to and fro at a temperature

between 300 and 500 C. This caused hot straining embrittlement, which once

more reduced the quality of the H.A.Z. at the tip concerned appreciably. A similar effect may occur at weld defects close to weld-crossings (fig. 11).

defect

first weld

FIG.11 DEFECT AT WELD CROSSING.

The transition temperature of the tránsverse (first) weld _{can be increased}

some 60°C, when the defect is at a critical distance from the (second)

longitudinal, weld.

The unfortunate dilemma for the structural engineer is, that _good

engineer-ing practice is a weldengineer-ing sequence as indicated in fig. 11. For then, there

is the smallest chance that weld defects _{occur. But whenever they occur,}

the situation may be far more dangerous than_{. when defeçts are present in a}

last made transverse weld.

5-17

Literature.

-.11/ J.J.W. Nibbering & A.W. Lafleinan.

"Low-cycle fatigue problems in shipbuilding; crack propagation in

coarse-grained.zones of thick plates".

Proc. Conference on fatigue of welded structures, Brighton, July 1970, paper 16.

/2/ H.C. van Elst.

"Over het onder'linge verband tussen _{brosse breukproeven met kleine} en grote proefstukken".

Lastechniek no. 8, 1967.

/3/ H. Kihara.

"Recent studies in Japan _{on brittle fracture of welded steel".}

11W-doc. X-291-61.

/'/ "A comparison of transition _{temperatures determined by small and large} scale tests on five steels".

Adzn. advisory committee on structurai _{steel. Report P2 - 1960.}

/5/ J.J..W. Nibbering & A.W. Lalleman.

"Low-cycle fatigue tests at low temperature with E.G.-welded 34 mm

plates of st. 52 Nb". 11W-doc. X-593-70.

/6/ JJ.W. Nibbering.

"Comparison between static C.0.D.-tests _{and Niblinic drop weight tests".} 11W-doc. W.G. 2912-168-72.

/7/. AA. Wells.

"Application of fracture mechanics at and beyond general yielding", British Welding Journal, Nov. 1963.