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Features of the brittle fracture of weld metaldeposited using electrodes with basic type coatingsI K Pokhodnya a , A O Korsun a , Yu Ya Meshkov b , G A Pakharenko b & A V Shevchenko ba E O Paton Welding Institute, Ukraine SSR Academy of Sciences ,b Ukr. SSR Academy of Sciences Institute of Metal Physics ,Published online: 04 Jan 2010.

To cite this article: I K Pokhodnya , A O Korsun , Yu Ya Meshkov , G A Pakharenko & A V Shevchenko (1987) Features of thebrittle fracture of weld metal deposited using electrodes with basic type coatings, Welding International, 1:7, 606-609, DOI:10.1080/09507118709453005

To link to this article: http://dx.doi.org/10.1080/09507118709453005

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A,fornuticheskup Svurku 1986 39 ( 10) 1-5

Features of the brittle fracture of weld metal deposited using electrodes with basic type coatings I K Pokhodnya and A 0 Korsun

Yu Ya Meshkov, G A Pakharenko and A V Shevchenko

E 0 Paton Welding Institute, Ukraine SSR Academy of Sciences

Ukr. SSR Academy of Sciences Institute of Metal Physics

Selected from Anomorirhrsskuyu S w k o 1986 39 (10); Reference AS/86/10/1; Translation 175

The difficulties of making welded joints which will be reliable in northern areas are in many cases due to the absence of reliable criteria for evaluating the conditions under which brittle fracture takes place. Modem fracture theory, based on the microscopic cleavage model,' makes it possible to link quantitatively particular structural components made of a particular metal with minimum brittle fracture stresses in the ductile-brittle transition temperature range (Fig.1) by simple equations: for carbon steels and armco iron (for which the stress intensity factor at the moment of fracture K, = 6MPa x m'/2)

R,, = Kf,d-'j2 ...[ 11

and for steel with globular shaped second phase particles

R,, = K f r d - r . . . [2]

where K, = 0.8MPa X rn'l2; is the microscopic cleavage stress, MPa; d is the grain size, nun; d, is the diameter of a globular second phase particle, mm. Since under conditions of overall macroscopic flow (when the applied stress has reached the yield stress) sub- microscopic cracks appear at the grain boundaries in metal, and since cleavage (shearing) of the second phase particles takes place, the level of resistance to brittle fracture is governed by the structural parameter initiating the larger new sub-microscopic crack.'

The ratio of the microscopic cleavage stress Kc to the nominal yield stress called the coefficient of

1 Relationship of mechanical properties of weld metal containing I % of nickel to temperature: S,, - true breaking stress; RZf - experimentally found microcleavage stress.

toughness K, in Ref.1, represents the potential capacity of a metal for resisting brittle fracture: the greater the difference between the values of uo.2 and K,, the greater the degree of over-stressing (caused by the deformation rate, test temperature or type of stress state) it can withstand without risking microscopic cleavages turning into serious fracture.

Our work consisted of making a study of whether a quantitative relationship can be established between the structural state of a weld metal and the particulars of its brittle strength and toughness.

Table 1 Chemical composition of the deposited metals, O h

Electrode Proportion by, c . Mn Si Ni S P mass of nickel in coating, yo

NO 0 N1 2 N2 4

~ -

0.095 0.95 0.36 - 0.019 0.024 0.090 1.05 0.28 1 .o 0.014 0.022 0.090 1.03 0.30 1.8-2.0 0.016 0.021

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2 Structures of weld metals (X2000, scale reduction 2/31 containing different amounts of nickel: (a) 0%; (b) 1%; (c) 2%.

Table 2 Mechanical properties of the metals investigated Electrode t,,, "C Not heat treated Heat treated

Reap d,,mm ~ ~ , ~ , m m R','!,MPa R$r,MPa, 1,% K, KCV, J/cm2 d,, mm mc > calculated M Pa .usinall1

NO + 20 - 40 - 60

N1 + 20 - 40 - 60

N2 + 20 - 40 - 60

0.010 390 400 410

0.008 370 390 410

0.006 430 490 520

1000 1000 1000

950 950 950

1000 1000 1000

1800 1800 1800

2010 2010 2010

2250 2250 2250

0.20 0.20 0.20

0.19 0.19 0.19

0.17 0.17 0.17

2.56 2.50 2.44

2.57 2.44 2.32

2.33 2.04 1.92

195-210 45-47 12-16

165-170 85-95 48-58

160-180 78-94 64-82

0.005 0.005 0.005

0.004 0.004 0.004

0.004 0.004 0.004

1100 1100 1100

1120 1120 1120

1140 1140 1140

Comment. In every cased,"= = 0.0008-0.001mm; R*:F. calculated using[3], is 111OMPa.

3 Distribution of non-metallic inclusions in the structure of a weld (X5400, scale reduction 314).

4 Fracture nuclei caused by non-metallic inclusion 1x2000, scale reduction 213).

The experiments were made on testpieces cut from the metal in the normalised zone of a multipass welded joint made using 4mm diameter ANO-26 electrodes with a basic type coating. Grade PNE-1 nickel powder (GOST 9722-79), its proportion by mass 0-4%, was added to the coating. The welds were made with AC from an STSh-500/80 transformer. The welding conditions were: I, = 180A, V,,, = 23-24V. The parent

metal was 14mm thick 09G2S steel (GOST 19282-73). The edges were prepared by variant B from GOST

The chemical composition of the metals deposited with the experimental electrodes is given in Table 1. The structures and fractures of the testpieces were studied with a JSM-35CF scanning electron microscope (JEOL, Japan). The composition of the non-metallic inclusions was determined with a Link-860 energy dispersion spectrometer (Link, Great Britain). The tensile tests were made at between +20 and - 196C with a UMM- 5 machine.

In addition to this the quality of the weld metal and its susceptibility to embrittlement were evaluated from the results of impact tests on testpieces with an acute angled notch (type IX, GOST 6996-66) at temperatures between + 20 and - 60C. The proportion f by volume and dimensions d,, of the non-metallic inclusions were detennined with a Quantimet-720 quantitative television microscope (Metals Research, Great Britain). The non- metallic inclusions smaller than 1 pm were also analysed with the JSM-35CF scanning electron microscope.

The results of the analysis with the electron microscope confirmed that the secondary structures of the testpieces investigated consisted of equiaxed ferrite grains with a fringe of bainite precipitates. A large number of non-metallic inclusions arranged both within the grain bodies and around their boundaries were also found. Special etching (using heated picric acid), camed out to reveal the grain boundaries (Fig.2), made it possible to determine the dimensions d, of the actual structural elements at whose boundaries sub-

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microcracks initiating microscopic cleavage form under general yield conditions.

Alloying weld metal with nickel reduced the grain size in it (Fig.2). As would be expected the mechanical properties of the metal changed. For instance, when the weld metal contained between 0 and 2% of nickel the impact strength of testpieces with an acute angled notch at -60C increased by about 400%, and the strength also increased slightly (Table 2).

The results of determining brittle strength were however unexpected. It was established that, in spite of the decrease in grain size d,, the microcleavage stresses kc remained practically unchanged, while according to calculations made using (11 there should have been a substantial increase in brittle strength R*:: (Table 2). The only reason why the calculated and experimental data did not agree can be that, in the testpieces investigated, microcleavages were caused not by the 'grain' source of sub-microcracks but by a different source forming large sub-microcrack nuclei which consequently developed under lower stresses. The non- metallic inclusions invariably present in the weld metal (Fig.3) might be this additional source of microcleavage in the testpieces studied. The decisive part played by the non-metallic inclusions in the fracturing of the weld metal was confirmed by the results obtained in Ref.2. Fractographic analysis of the surfaces of testpieces which had failed by brittle fracture revealed the

42 0,s 1,O 1,4 48 5 2 dine ~m

5 Histogram of distribution of non-metallic inclusions by size (weld metal containing 1 % of nickel).

6 Structure of weld metal containing no nickel after heat treatment (X2400, scale reduction 314).

existence of fracture nuclei caused by non-metallic inclusions (Fig.4).

It was proved in Ref.3 and 4 that, under certain conditions, the level of brittle strength depends on the dispersion and distribution of the globular second phase particles. At the same time, with a proportion by volume of less than 4.5%, the particles within the bodies of grains cannot be fracture sources. In this case fracture is initiated by second phase particles at the grain boundaries, and the microcleavage stress is found using the equation

R,, = K,d- r . . . [3]

where K , = 1.lMPa X m1'2. The most likely diameter dinc of these particles in the testpieces investigated (Table 2) was found using the histogram (Fig.5) of distribution of non-metallic inclusions by dimensions.

Table 2 also contains the results of calculating the microcleavage stress R'i:r using (31. Comparison of the calculated' and experimental (RZZ) data leads to the conclusion that the non-metallic inclusions play the decisive part in the brittle fracture of the steel investigated.

In order to confirm the law which had been established an additional experiment was carried out. Testpieces were quenched (austenitisation temperature 9SOOC) and then tempered at 380C. This heat treatment caused a different structure to form (Figd), different as to both phase composition and grain size, but had no effect on the dimensions of the non-metallic inclusions, since these are complex alloyed oxides (Fig.7), whose temperature of solution is, according to Ref.5, - 160O"C, well above the heat treatment temperature. It can be seen from Table 2 that the heat treatment also scarcely affected the microcleavage stress R ~ i Y obtained by calculation using [3].

It can thus be argued that the level of brittle strength of the testpieces considered is governed not by the grain size but by the dispersion of the non-metallic inclusions. Moreover the coefficient of toughness for the steel investigated, K, = ( R J U ~ , ~ ) also depends to a great extent on the size of the non-metallic inclusions (Table 2). Decrease in grain size increases the nominal yield stress u ~ , ~ , but with a fixed and constant level of microcleavage stress kc this is inevitably accompanied by a drop in toughness margin. The greater the rate at which the yield stress increases the greater will be this drop.

It should therefore be argued that the beneficial effect associated with the presence of nickel in weld metal has not essentially been realised (only the impact strength has been increased). Weld metal containing 2% of nickel has a lower coefficient of toughness (at temperatures between 20 and - 60C) than weld metal containing no nickel. Using the model of microcleavage for analysing the relationship of the resistance to brittle fracture of weld metal to the structure has made it possible to establish the effective structural parameter controlling the process of brittle fracture. This

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7 Alloying element contents of an - lpm diameter non-metallic inclusion. E, keV parameter proved, in the case of the steel investigated, to be the size of the non-metallic inclusions. This is why alloying with nickel, which does not affect the dispersion of these inclusions, does not, unless the system of deoxidation is correctly selected, increase the brittle strength of weld metal. The increase in impact strength accompanying increase in nickel content does not represent the true toughness margin of the metal, and is due to increase in the yield stress. Reference 6 also found that the beneficial effect of nickel (with manganese contents of over 1%) on the mechanical properties of welds is reduced.

Establishment of the relationship between the strength and ductility of weld metal and the structural parameters makes it possible to find ways of further improving the mechanical properties, consisting of reducing the size of the structure elements capable of initiating sub-microcracks. The research described in Ref.4 proved that the combination of brittle strength and toughness margin is optimal with a ratio of 25:l between the grain size in the steel and diameter of a globular particle of the second phase, with the maximum dispersion. It was for instance proved in Ref.7 and 8 that, when the Mn/Si ratio is increased, the size of the non-metallic inclusions is reduced.

Therefore, in order to provide welded joints in steel with high service particulars at subzero temperatures, the systems by which the weld metal is alloyed and deoxidised must be selected in conjunction.

CONCLUSIONS 1 Globular shaped non-metallic inclusions, which are

products of the deoxidation of weld metal with manganese, silicon and titanium, may be sources of brittle fracture, thus limiting the possibility of increasing the toughness margin of a weld. 2 The existence of these non-metallic inclusions prevents advantage from being taken of the beneficial effect of nickel on brittle strength (in spite of decrease in stnictural grain size and increase in impact strength of weld metal). 3 One way of improving the brittle strength and toughness of weld metal is to match the reduction in size of the grains and non-metallic inclusions.

References 1 Meshkov Yu Ya: T h e physical fundamentals of the fracturing of steel structures.' Pub1 Naukova duma, Kiev, 1981. 2 bkhodnya I K el a,! T h e failure of single pass meld metal at temperatures within the brittleductile transition range.' Art. Svurku

3 Meshkov Yu Ya er oL 'Breakdown of steel containing granular cementite.' Aferollofuiko 1983 5 (3) 94-97. 4 hleshkov Yu Ya er ut The brittle fracture of carbon steels mith different granular cementite distributions.' Ukr. SSR Academy of Sciences Institute of Metal Physics 1984 13. hl/s deposited at VlNITI on 21.12.84 No.8217-84.

5 Podgaetskii V V: "on-metallic inclusions in welds.' Pub1 hlashgiz. Moscow, 1962. 6 Taylor D S: 'Development of MMA electrodes for offshore fabrication.' IVeId. 1.1983 15 f8) 438-443.

7 bkhodnya I K ef ot T h e non-metallic inclusions in welds made using electrodes with mi le and ilmenite coatings.' A,% Svorko 1976 (9) 8-11.

8 Widgety D J: 'Deoxidation practice for mild steel weld metal.'

1979(3) 1-4.

\\kldJ. 1976 55 (3) 125-137.

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