POSSIBILITIES AND LIMITATIONS TO IMPROVE THE WELDABILITY ...

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POSSIBILITIES AND LIMITATIONS TO IMPROVE THE WELDABILITY OF LOW CARBON 12Cr FERRITIC STAINLESS STEEL FOR EXPANDED INDUSTRIAL APPLICATIONS


E. Deleu

A. Dhooge

E. Taban

E. Kaluç

ABSTRACT Ferritic stainless steel X2CrNi12 (EN 10088) is generally appreciated for its relatively low cost and good resistance to wet abrasion and mild environments, but unfortunately its weldability is restricted. Typical applications up to now include railway wagons for coal and iron ore, mining and mineral process and transport equipment, bus frames and chassis, silos, etc. This steel grade nowadays can be fabricated cost effectively with low carbon and impurity levels appreciably improving both the weldability and mechanical properties. In this case, long-term maintenance costs of assemblies produced using this ‘clean’ X2CrNi12 stainless steel will be low, with a suitable coating providing sufficient protection for several decades. For other applications, the use of weldable X2CrNi12 is also more economical than higher alloyed stainless steels. Moreover, joining the steel by laser welding without filler metal should be considered. The present paper provides an overview of results of a research project initiated by the Belgian Welding Institute investigating the possibilities of this modified low carbon ferritic stainless steel. The main objective of this work was to demonstrate that it offers great application ranges for constructing purposes which should expand its application field substantially. In the near future broadening this steel family with even higher mechanical properties allowing further reduction of plate or wall thickness and of production costs will be explored. IIW-Thesaurus keywords: COD; Corrosion; Ferritic stainless steels; Impact toughness; Filler materials; Mechanical properties; Reference lists; Simulating; Stainless steels; Steels; Structural steels; Toughness; Weldability; Welding.

Mr. Eddy DELEU ([email protected]), Senior Research Engineer is with the Belgian Welding Institute (BWI), Gent (Belgium). Prof. Alfred DHOOGE (alfred. [email protected]) is with the Mechanical Engineering Department, Faculty of Applied Science, University of Gent, Gent (Belgium). Dr. Emel TABAN (emelt@kocaeli. edu.tr) and Prof. Dr. Erdinç KALUÇ (ekaluc@kocaeli. edu.tr) are with the Mechanical Engineering Department, Engineering Faculty, Kocaeli University, Kocaeli (Turkey). Doc. IIW-1918-08 (ex-doc. IX-2228r1-07) recommended for publication by Commission IX “Behaviour of metals subjected to welding”. Welding in the World, Vol. 53, n° 9/10, 2009 – Peer-reviewed Section

INTRODUCTION Ferritic stainless steels are the second most widely used group of stainless steels, due to their good corrosion resistance and lower production costs, as compared with austenitic stainless steels. They have been considered in the past as exhibiting low weldability and have therefore mostly been used for applications that do not require welding, since a fully ferritic microstructure has rather poor toughness at low temperature and inferior high temperature strength compared to austenite. For these reasons, in some predominantly ferritic steels, a small amount of austenite forms at high temperatures and may transform to martensite on cooling, to develop


12 %Cr ferritic/martensitic steels, with close control of the carbon content and ferrite/martensite balance, to avoid the extremes of completely ferritic or martensitic microstructures. The first generation of 12 %Cr steels is known as 3Cr12 stainless steels which were developed in the late 1970s with 0,03 %C. It is still produced nowadays and conforms to grade 1.4003 or X2CrNi12 of EN 10088 “Stainless Steels” and of EN 10028-7 “Flat products made of steels for pressure purposes” and to UNS S41003 of ASTM A240 “Standard Specification for Chromium and Chromium-Nickel Stainless Steel Plate, Sheet and Strip for Pressure Vessels and for General Applications”. A series of investigations describing the research and use of 3Cr12 steel in various applications can be found in literature [1-6]. Although 3Cr12 has excellent corrosion resistance in many environments, its weldability is limited. In the early 1990s, the second generation, namely 5Cr12, was developed to give better toughness, but weldability remained limited due to the higher carbon content [7-8]. Initial applications of these steels were in materials handling equipment in corrosive/abrasive environments, but they were later used extensively in the coal mining industry, for cane and beet sugar processing equipment and for bulk transport of coal and gold, road and rail transport, power generation and in aerospace engineering. Contemporary steel manufacturing plants facilitate the production of X2CrNi12 stainless steel with low carbon (< 0,015 %) and impurity levels to improve weldability, thus displaying both the advantages of stainless steels for corrosion resistance and engineering properties of carbon steels [1, 9-12]. For long-term maintenance costs, this modified stainless steel requires few coating renewals, offering substantial economic and considerable environmental advantages. For various applications, the use of this steel with improved weldability would be more economical, compared to higher alloyed stainless steels [11, 13]. Matching welding electrodes are commercially available for welding of X2CrNi12 stainless steel. However, it is not recommended in applications where impact, fatigue or any other form of non-static loading is anticipated. Reported weldability studies have shown instead that austenitic stainless steel consumables can be used to produce welds, with a minimum risk for heat-affected zone (HAZ) hydrogen cracking because of the high solubility and low diffusivity of hydrogen in these weld metals, preventing hydrogen entering the adjacent HAZ and to ensure deposition of tough weld metal with adequate properties required for structural purposes [1, 14-16]. In the present paper, the outcome is summarized of a research project on this low carbon X2CrNi12 stainless steel executed for a group of steels and consumables suppliers, manufacturers and end users. Different welding processes and various austenitic and duplex stainless steel consumables were applied and

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the respective joints were evaluated, based on their microstructural, mechanical, toughness and corrosion properties.

RESEARCH PROGRAMME The main objective of the research project was to evaluate the applicability of this modified, low carbon, ferritic stainless steel, by examining its possibilities for use in structural applications like pressure vessels, beams, pipelines, bridges or parts thereof, as alternatives for non-alloy structural steels or other types of stainless steel. Depending on the interest of the project members, plates with thicknesses of 6 mm, 12 mm, 20 mm and 30 mm were investigated, while dissimilar welds made between X2CrNi12 stainless steel and S355J2 (EN 10025) structural steel were incorporated. Applied welding processes were manual or shielded metal arc welding (SMAW), semi-automatic welding such as gas metal arc welding (GMAW) or flux cored arc welding (FCAW), but also submerged arc (SAW) and laser welding, the latter without filler metal. Prior to actual welding, thermal weld simulations were executed to investigate the suitability of the test material, by applying single or double thermal cycles with varying peak temperatures and cooling rates that are similar to those effectively realized during welding.

X2CrNi12 parent metal properties The parent metal microstructure in as-delivered condition, i.e. tempered at 690 °C, mainly consists of ferrite (more than 60 %) with appreciable amounts of tempered martensite for increased tensile properties, see Figure 1. Because the main objective of this research project was to evaluate the suitability of this weldable X2CrNi12 stainless steel for structural applications, guaranteed yield and tensile strengths were compared to those specified for a typical structural steel, such as S355J2. The chemical composition of the heats and the mechanical properties of the relevant plates are given in Tables 1 and 2, respectively. The chemical analyses of all heats conformed to that specified in EN 10088, although the project target content for silicon and phosphorus was 0,45 % and 200 ppm, respectively. Except for some UTS values, the mechanical properties of the plate material fulfilled the requirements of S355J2. During the project it became clear that producing 12 %Cr stainless steel combining very attractive properties is not readily accomplished, as the latter may be affected by small deviations in tempering treatment, or by small additions of elements like boron (above 20 ppm or 30 ppm). Also notch impact toughness may vary within the same plate. Welding in the World, Vol. 53, n° 9/10, 2009 – Peer-reviewed Section

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a) As-delivered parent metal

b) Simulated at 1 350 ºC

c) At 1 000 °C

d) At 800 °C

e) At 1 350 °C followed by a second peak at 1 000 °C

f) At 600 °C

Figure 1 – Parent metal and weld-simulated microstructures observed after etching in Vilella’s reagent for the 12 mm-thick X2CrNi12 plate 889879, all valid for low heat input welding Table 1 – Chemical composition (in weight %, except otherwise specified) of the heats used for producing the ferritic stainless steel plates (data from steel supplier) Heat 51351 a 25341 b 56671 c 58841 d EN 10088 a b

C 0,016 0,014 0,012 0,010 < 0,030

Si 0,25 0,45 0,26 0,32 < 1,00

Mn 0,96 0,99 0,95 0,97 < 1,50

P (ppm) 230 210 350 320 < 400

Heat used for the 6 mm-thick plates for actual welding. Heat used for the 12 mm-thick plate for weldability test programme (thermal weld simulation).

Data shown in italics are specifications according to X2CrNi12 (EN 10088). Welding in the World, Vol. 53, n° 9/10, 2009 – Peer-reviewed Section

S (ppm) 5 10 10 11 < 150 c d

Cr 12,45 12,32 12,45 12,38 10,5-12,5

Ni 0,42 0,44 0,51 0,71 0,3-1,0

N (ppm) 80 61 80 101 < 300

Heat used for the 12 mm and 20 mm-thick plates for actual welding. Heat used for the 30 mm-thick plates for actual welding.


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Table 2 – Mechanical properties of the ferritic stainless steel plates (data from steel supplier) Plate 229975 889879 252276 252261 263902

a b c d e

EN 10025

Yield strength Tensile strength (MPa) (MPa) 363 ; 359 527 ; 525 497 633 363 ; 362 502 ; 500 352 ; 353 507 ; 506 450 ; 437 591 ; 582 ≥ 355 (≤ 16 mm) 510-680 ≥ 345 (16-40 mm)

Strain at fracture (%) 31 ; 33 37 30 ; 32 28 ; 29 25 ; 24

Hardness (HB) 146 172 129 136 160

Impact toughness at -20 °C (J [*] or J/cm2 [**] ) 63-67-70 / 67 ; 83-68-67 / 73 [**] Not mentioned 101-108-123 / 111 ; 294-183-75 / 184 [*] 73-60-69 / 67 ; 96-88-76 / 87 [*] 150-81-71 / 101 ; 111-77-80 / 89 [*]

≥ 18

-

≥ 27 (mean) ; ≥ 20 (individual) [*]

a

6 mm-thick plate used for actual welding and originating from heat 51351. 12 mm-thick plate used for weldability test programme and originating from heat 25341. 12 mm-thick plate used for actual welding and originating from heat 56671. d 20 mm-thick plate used for actual welding and originating from heat 56671. e 30 mm-thick plate used for actual welding and originating from heat 58841. Data shown in italics are specifications according to S355J2 (EN10025) structural steel. Values separated by symbol “;” originate from two different positions within the plate. b c

Weldability of low carbon X2CrNi12 ferritic stainless steel

temperature and so to detect microstructural transformations during heating and cooling of the simulated heat-affected zone.

Thermal weld simulations were executed to assess the influence of single and double cycles on base metal morphology, hardness and impact toughness, by creating various synthetic HAZ microstructures. To achieve this, the test material was subjected to various peak temperatures between 800 °C and 1350 °C, with a holding time of only 0,1 s and to cooling rates of 50 °C/s (low heat input) or 15 °C/s (high heat input). Double thermal cycles, simulating the effect of multiple welding passes, were applied using a first peak temperature of 1 350 °C, followed by a second peak temperature of 1 000 °C or 600 °C and with the same holding time and cooling rates as used for single simulations.

Specimens of 55 mm x 10 mm x 10 mm were extracted, generally with their length transverse to the rolling direction, while no additional heat treatment after weld simulation (or PWHT) was applied, as this is normally not done in practice when using austenitic types of filler metal demanding a too high solution annealing temperatures.

In order to extrapolate these conditions to a real welding situation, corresponding heat inputs for the range of plate thicknesses used within the project are given in Table 3, according to calculations proposed by Rykalin. Dilatometric measurements were done to follow the thermal expansion of the specimen thickness with

Table 3 – Heat inputs invoking selected cooling rates for the investigated plate thicknesses valid for X2CrNi12 stainless steel Plate thickness (mm) 6 12 20 30

Cooling time [cooling rate] between 800 °C and 500 °C 6 s [50 °C/s] 20 s [15 °C/s] 6 s [50 °C/s] 20 s [15 °C/s] 6 s [50 °C/s] 20 s [15 °C/s] 6 s [50 °C/s] 20 s [15 °C/s]

Heat input (kJ/mm) Preheat temp. of 20 °C

Preheat temp. of 150 °C

0,31 0,57 0,62 1,14 1,04 1,90 1,27 2,85

0,21 0,39 0,43 0,78 0,71 1,30 0,77 1,94

Heat inputs given in italics correspond with a so-called three-dimensional cooling regime, all others with a two-dimensional cooling regime.

One specimen per combination of peak temperature and cooling rate was used for microstructural analysis and Vickers hardness measurements using a load of 49 N (HV5). Three other samples that had been subjected to the same thermal cycle were through-thickness notched to determine the impact toughness at a temperature of –20 °C. Unaffected base metal specimens were also investigated in the same way, the results of which may differ from those mentioned before. Indeed the base metal hardness for plate 889879 was about 205 HV5 irrespective of specimen orientation. Impact toughness measured at –20 °C instead strongly depended on orientation and changed for transverse and longitudinal directions, respectively from 141 J to 224 J. Moreover, fairly large scatter between data was noticed for each condition, with maximum ratios between individual values of about two. Also, during the welding part of the project, substantial differences between plates were observed. Simulated HAZ hardness From dilatometric records, it was clear that during heating at 100 °C/s, transformation of ferrite or martensite to austenite starts at about 850 °C and finishes at about 1 000 °C, while transformation to δ-ferrite starts at about 1 200 °C and finishes above 1 300 °C. During cooling from 1 350 °C at 15 °C/s, transformation to martensite starts at about 400 °C and finishes at about 300 °C. The highest hardness of 260 HV5 to 280 HV5 in the simulated HAZ was detected for peak temperatures Welding in the World, Vol. 53, n° 9/10, 2009 – Peer-reviewed Section

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between 900 °C and 1100 °C, see Figure 2. This is attributed to the presence of untempered martensite from the dual-phase morphology still retaining a fine microstructure. At peak temperatures of 1 350 °C, excessive growth of ferrite grains occurs which during cooling partially transforms to austenite and finally to martensite. For the lowest peak temperature of 800 °C, no transformation and therefore no refining occur.

All of this shows that HAZ hardness in the investigated low carbon ferritic stainless steel is mainly governed by grain size and to a lesser extent by the portions of tempered versus untempered martensite.

A second thermal cycle at 1 000 °C on HAZ microstructures adjacent to the fusion line only increased hardness by some 20 HV5, see also Figure 2, probably because of the creation of new untempered martensite with only little refining effect. Applying a second cycle at 600 °C instead softened the same coarse grained HAZ microstructures by about 25 HV5 due to martensite tempering. No effect was noticed for a similar second cycle on the hardest but fine HAZ microstructures created after a simulation at a first peak temperature of 1 000 °C.

HAZ impact toughness at –20°C was at least as high as that of the base metal for peak temperatures below 1 100 °C, see Figure 3 and within this temperature range the influence of sample orientation was still obvious. For higher peak temperatures and thus for HAZ microstructures closer to the fusion line the impact toughness was severely reduced to about 10 J and the effect of sample orientation disappeared. For these regions also no distinction in toughness between the selected extreme cooling rates was noticed. Unfortunately a second thermal cycle with a peak temperature at 1 000 °C or 600 °C did not recover toughness as mean values still remained around 10 J, see again Figure 3.

Differences in hardness due to variations in cooling rate or in sample orientation in general were only marginal.

Simulated HAZ impact toughness

Figure 2 – Mean HV5 hardness of base metal and single or double cycled weld simulated microstructures of the X2CrNi12 ferritic stainless steel [L or H = low or high heat input; 1 000 or 600 = second peak temperature in ºC, if applicable]

Figure 3 – Mean impact toughness at -20 °C of base metal and single or double cycled weld simulated microstructures for the X2CrNi12 ferritic stainless steel [L or H = low or high heat input; 1 000 or 600 = second peak temperature in °C, if applicable] Welding in the World, Vol. 53, n° 9/10, 2009 – Peer-reviewed Section


Apparently the main factor yielding disappointing HAZ properties for this ferritic stainless steel is the grain size which is excessive when the material is exposed to temperatures as high as 1 350 °C. Due to its nature, grain refinement is not readily achieved and so multi-pass welding is no cure against low HAZ toughness. Photomicrographs showing parent metal and weld simulated microstructures after etching in Vilella’s reagent are given in Figure 1 illustrating the above observations. Whether or not the present conclusions should give rise to doubts concerning the material’s weldability was clarified during evaluation of real weldments. It is recognized that weld simulation normally produces (sometimes extremely) conservative data as larger areas with identical microstructures are created than in real welds, where local zones might be protected by surrounding regions with superior properties.

Selection of filler metals for X2CrNi12 ferritic stainless steel Apart from the consideration of achieving adequate corrosion resistance when selecting filler metals for stainless steels, the prevention of cracking should also be accounted for. Hot cracking is of most concern in austenitic stainless steels, while cold cracking may occur in ferritic stainless steel weldments that are embrittled by grain coarsening and/or second

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phase particles. In many cases, ferritic filler metals are replaced by austenitic filler metals that include a small amount of ferrite to avoid the occurrence of both types of cracking. Although in this case weld metal grain growth normally is not a problem, grain growth at the HAZ should not be overlooked while discrepancies in thermal expansion may cause difficulties. Also the low yield strength of austenitic stainless steel weld metals may cause undermatching conditions in terms of tensile properties. For the present project, mainly austenitic filler metals were selected such as AWS E309L, E309LSi, ER309L or E309LT containing large amounts of chromium and nickel, lower alloyed ER308LSi and molybdenum alloyed ER316L, ER316LSi or E316LT for improved pitting corrosion resistance. One weld was also made with a duplex stainless steel filler metal (ER2209) that should not create yield strength undermatching problems. Typical chemical compositions of the filler metals used are given in Table 4.

Welding of low carbon X2CrNi12 ferritic stainless steel A summary of the combinations of welds investigated is given in Table 5 mentioning also the selected types of consumables. The respective codes represent the identifications proposed during the execution of the entire test programme. In total 16 welds, each with a

Table 4 – Chemical composition of the consumables used in this research work (according to AWS) Type of consumable E309L-XX ER308LSi ER309L ER309LSi ER316L ER316LSi ER2209 E309LTX-X E316LTX-X

C

Mn

P

S

Si

Cr

Ni

Mo

N

Cu

0,04 0,03 0,03 0,03 0,03 0,03 0,03 0,04 0,04

0,5-2,5 1,0-2,5 1,0-2,5 1,0-2,5 1,0-2,5 1,0-2,5 0,5-2,0 0,5-2,5 0,5-2,5

0,04 0,03 0,03 0,03 0,03 0,03 0,03 0,04 0,04

0,03 0,03 0,03 0,03 0,03 0,03 0,03 0,03 0,03

0,90 0,65-1,00 0,30-0,65 0,65-1,00 0,30-0,65 0,65-1,00 0,90 1,00 1,00

22,0-25,0 19,5-22,0 23,0-25,0 23,0-25,0 18,0-20,0 18,0-20,0 21,5-23,5 22,0-25,0 17,0-20,0

12,0-14,0 9,0-11,0 12,0-14,0 12,0-14,0 11,0-14,0 11,0-14,0 7,5-9,5 12,0-14,0 11,0-14,0

0,75 0,75 0,75 0,75 2,0-3,0 2,0-3,0 2,5-3,5 0,5 2,0-3,0

0,08-0,20 -

0,75 0,75 0,75 0,75 0,75 0,75 0,75 0,50 0,50

Single values shown are maximum percentages.

Table 5 – Complete welding programme realized during the project Welding process

6 mm 12Cr

Manual

-

Semi-automatic

R9 * [309]

Submerged arc

-

Laser

-

12 mm 12Cr V9 [309] B9 ; B8 ; B6 * [309 ; 308 ; 316] E9 ; E6 [309 ; 316] V [-]

20 mm 12Cr

30 mm 12Cr

12 mm 12Cr-S355J2

20 mm 12Cr-S355J2 M9 [309]

-

-

-

F9 ** [309] A9 [309]

K9 ** [309] P9 ; PD [309 ; Duplex]

U9 ** [309]

-

-

C9 [309]

-

-

-

-

* = GMAW with Ar/He/2-3CO2 gas protection. ** = FCAW with Ar/12-18CO2 gas protection. Welding in the World, Vol. 53, n° 9/10, 2009 – Peer-reviewed Section

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total length of 2 m, were realised in plates between 6 mm and 30 mm-thick and by using different welding processes. Dissimilar welds between X2CrNi12 stainless steel and structural steel S355J2 were also included since such types of joints are often applied in practice for attaching accessories to the main part of the construction. Although the extent of experiments depended on the weld combination, most of the joints were evaluated on the basis of their appearance, weld shape and position of welding passes, weld metal chemical composition, hardness, tensile and bend properties, notch impact and fracture toughness at various temperatures down to –60 °C, fatigue properties and resistance against atmospheric attack. Because of the general character of the present document and of the large extent of the complete test programme, only few detailed results are given so that the following paragraphs describe an overview of the most pertinent observations. Practically all of these also apply for the dissimilar welds except for the comments concerning HAZ toughness, which strictly speaking are only valid for the X2CrNi12-side. Metallography Metallographic examinations showed that sound welds can be produced in low carbon X2CrNi12 ferritic stainless steel, if proper welding parameters are applied. As expected from weld simulation work, excessive grain growth was observed at the heat-affected zone, which mainly consisted of ferrite including a lot of islands containing martensite. Evidence of this is given in the photomacrograph of Figure 4 where the difference in HAZ grain growth between both sides is striking and in the photomicrograph of Figure 5. The latter also illustrates the occurrence, in some cases, of precipitates located inside the large ferrite grains. Observations at the heat-affected zone of all actual welds were very similar to those given in Figure 1 and relevant comments given earlier. Both, lack of side wall fusion and lack of penetration occurred in some of the welds, which

Figure 4 – Photomacrograph after etching in Vilella’s reagent of the 12 mm-thick dissimilar flux cored arc weld made with an E309LT-1 type of consumable (with S355J2 structural steel on the left and X2CrNi12 ferritic stainless steel on the right) Welding in the World, Vol. 53, n° 9/10, 2009 – Peer-reviewed Section

might escape radiographic detection. Maximum HAZ hardness of 270 HV5 to 300 HV5 was noticed somewhat away from the fusion line and these observations agreed extremely well with those made during weld simulations, see before. Maximum weld metal hardness for welds made with austenitic or duplex stainless steel filler metals varied respectively from 200 HV5 to 240 HV5 and from 240 HV5 to 250 HV5. The hardness of the fused metal of the laser weld made without filler metal was limited to 270 HV5. Tensile properties Transverse tensile strength of all welds was above that of the actual base metal even in the presence of important lack of side wall fusion. It should be noted that tensile strength of the 12 mm and 20 mm-thick plate materials used for actual welding was about 500 MPa, which was the minimum project target strength. This means that the use of austenitic stainless steel filler metals represents a real risk for mechanical undermatching, certainly if parent metal tensile strength is beyond 600 MPa, like for the plate material that was produced for the weldability test programme, see Table 2. Moreover longitudinal weld metal tensile tests revealed that in these cases an ultimate tensile strength above 600 MPa is virtually impossible to achieve in practice. Obviously, welds made with duplex stainless steel filler metals bear no such risk at all. Ductility Room temperature ductility, as demonstrated by bend testing, was excellent. In a lot of cases, samples contained shallow undercuts prior to testing but these appeared to be largely harmless. One sample removed from the 30 mm-thick submerged arc duplex weld (weld PD of Table 5) fractured at a bending angle of about 45°. This was attributed to the weld shape, see Figure 6, where the eccentric position of shallow

Figure 5 – Photomicrograph after etching in Vilella’s reagent of the HAZ of the 12 mm-thick GMAW weld made in the X2CrNi12 ferritic stainless steel with an ER309LSi type of consumable


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this was attributed to the base metal properties, see earlier. Examples of notch impact toughness test data are summarized in Table 6 for three homogeneous butt welds made in 12 mm-thick plates. Some investigations were undertaken to examine the possible cause of low toughness at the heat-affected zones of most of the welds. Available macro sections of nearly all welds were examined at the heat-affected zone at various distances from the fusion line and at four thickness positions from sub-surface to mid-thickness while ASTM grain size numbers were measured. During this examination account was taken of the most probable microstructures that were sampled in the respective test specimens depending on their notch position.

Figure 6 – Photomacrograph after etching in Vilella’s reagent of the 30 mm-thick homogeneous submerged arc weld made in X2CrNi12 ferritic stainless steel with duplex filler metal (ER2209)

In general poor individual fusion line toughness corresponded with coarse grains (ASTM grain size number M10, 2 or 3). No grain coarsened HAZ microstructures were detected in the 30 mm-thick submerged arc weld made with an ER309L type of wire (weld P9 of Table 5) possessing indeed a low transition temperature or high toughness. On the other hand, the weld with the lowest transition temperature and thus with the best impact behaviour showed much coarser HAZ microstructures. No other direct correlations between fusion line or HAZ toughness and grain size could be detected.

capping passes or “wings” with long epitaxial grains directed perpendicular to the plate surface stimulated fracture occurrence. Such an extreme weld profile should therefore be avoided in practice although capping passes are sometimes deposited eccentrically for improved weld appearance.

Heat-affected zone CTOD fracture toughness at –20 °C, which was only determined for weldments with a thickness of 20 mm or higher, was disappointing as minimum values of 0,10 mm for sets of three identical tests were very difficult to realize.

Fracture toughness properties Low temperature notch impact toughness of the welds was promising if grain growth at the heat-affected zone could be limited. In such cases the 27 J transition temperature was situated between –20 °C and –50 °C for notch positions ranging from the weld metal centre to the heat-affected zone at 5 mm from the fusion line. In other cases, the transition temperature increased to 0 °C or higher, mainly due to low fusion line or heataffected zone toughness, and this observation was incompatible with the intended project objectives. This was also the case for the laser weld demonstrating quite low impact toughness. For some welds scatter of HAZ impact toughness was unexpectedly high and

Correlation between weld cooling rate and HAZ impact notch toughness Minimum and maximum cooling rates applied for seven different welds with a thickness of 12 mm or 20 mm were calculated based on the reported heat inputs, the plate thickness and the measured preheat and interpass temperatures and these were correlated with the respective impact properties at –20 °C. It was aimed to

Table 6 – Notch impact toughness measured on standard test samples removed from 12 mm-thick butt welds made in X2CrNi12 ferritic stainless steel Welding process / Type of consumable (acc. to AWS)

Notch position [***]

Mean impact toughness (J) at

Transition temperature [****]

+20 °C

0 °C

-20 °C

-40 °C

-60 °C

GMAW / ER309LSi [*] (weld B9 of Table 5)

WMC FL FL+2

142 48 209

128 40 61

122 26 65

Not tested

Not tested

–20 °C

GMAW / ER308LSi [*] (weld B8 of Table 5)

WMC FL FL+2 FL+5

138 70 145 293

125 18 68 236

122 37 65 40

Not tested

Not tested

0 °C to +10 °C

SAW / ER309L [**] (weld E9 of Table 5)

WMC FL FL+2

Not tested

Not tested

62 43 39

55 38 31

48 37 19

–50 °C

[*] Protection gas for GMAW was Ar/He/2-3CO2 [**] Flux for SAW was a basic non alloying agglomerated flux. [***] WMC = weld metal centre; FL = fusion line; FL+2 or FL+5 = HAZ at 2 mm or 5 mm from FL respectively. [****] Transition temperature was based on mean toughness of 27 J. Welding in the World, Vol. 53, n° 9/10, 2009 – Peer-reviewed Section

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define a window of extreme cooling rates guaranteeing proper HAZ toughness for X2CrNi12 ferritic stainless steel welds. Unfortunately no such trend was observed although accumulated cooling times for two or three consecutive weld passes were considered. Moreover, accumulated cooling times calculated relative to the plate thickness eventually combined with the number of weld passes still revealed no correlation between cooling conditions and remnant HAZ impact toughness. Fatigue behaviour Fatigue resistance, measured only for some 12 mmthick welds, including the laser weld with limited toughness, was excellent. Only the GMAW-weld B9 of Table 5 containing an intermittent lack of side wall fusion revealed fatigue behaviour inferior to that anticipated for a base metal with a tensile strength for instance of 600 MPa. Far better results were obtained for the submerged arc weld made with an ER309L type of wire (weld E9 of Table 5), see Figure 7. The purple and yellow lines in this figure represent the expected fatigue resistance for base metals with a tensile strength of 500 MPa and 600 MPa respectively. This and all other welds were tested with the excess of weld metal removed flush with the parent metal and with a fatigue ratio R of 0,1. Environmental exposure Finally, the endurance against atmospheric attack determined during salt spray testing of coated welds made in low carbon X2CrNi12 ferritic stainless steel was also very promising even when tested under severe conditions created by an artificial scratch across the weld. Again as expected, the welds made with the high alloyed filler metals demonstrated a higher resistance than those made with ER308LSi type of austenitic filler metals. Under normal atmospheric conditions though, all welds offered the possibility to prevent further deve-

lopment of corrosion products, once initiated. Typically welds with extremely good mechanical properties were classified as somewhat less resistant against atmospheric attack.

Effect of filler metal Weld metal hardness for welds made with ER308LSi consumables was 10 HV5 to 20 HV5 higher than that of welds made with E(R)309L or E(R)316L types of consumables. This was confirmed by the higher weld metal tensile strength of the former type of filler metal. On the other hand yield strength was lower for ER308LSi than for E(R)309L types of wire and the former revealed a better ductility during tensile testing. The highest yield strength was obtained with E(R)316L types of consumables though accompanied by a slight decrease in ductility. Weld metal notch impact toughness at –20 °C was very similar for all three austenitic filler metals while fusion line toughness, where still a certain amount of weld metal was involved, was lower for the reference consumable than in case of the other filler metals. Resistance against atmospheric attack was definitely affected by the type of consumable in protected condition and artificially damaged. In these cases E(R)316L types of filler metal often improved the resistance of the whole system with regard to E(R)309L types of filler metal while ER308LSi consumables, due to their lower alloying, demonstrated an inferior behaviour. So the use of a more economical consumable like ER308LSi compared with a higher alloyed E(R)309L types results in a small concession concerning yield strength (for GMAW-welding about 325 MPa compared with about 345 MPa). From this point of view, it is certainly worthwhile to consider using the cheaper type of consumable for many applications of X2CrNi12 ferritic stainless steel under less aggressive circumstances. The highest tensile properties are achieved by using

Figure 7 – Fatigue resistance of the 12 mm-thick submerged arc weld made in X2CrNi12 ferritic stainless steel with an ER390L type of consumable Welding in the World, Vol. 53, n° 9/10, 2009 – Peer-reviewed Section


E(R)316L types of consumables, which should also yield improved local corrosion behaviour.

INTERPRETATION AND CONCLUSIONS The research work on X2CrNi12 ferritic stainless steel with improved weldability executed for a group of members has permitted the following conclusions to be drawn: 1. Manufacturing 12Cr ferritic stainless steel conforming to X2CrNi12 (EN 10088) but with reduced amounts of carbon and impurities is possible with reasonable production costs so that the material possesses tensile and toughness properties complying with requirements specified in EN 10025 for non alloy structural steel type S355J2. 2. The weldability determined through systematic research based on thermal weld simulations is rather limited although the severe conservatism inherent to such an approach should be accounted for as could be demonstrated by some of the real welds. 3. In general, high productivity and sound homogeneous and dissimilar welds can be made by shielded metal arc, gas metal arc, submerged arc and flux cored arc welding using austenitic or duplex stainless steel consumables commercially available today. Hardness at the heat-affected zone of this stainless steel can easily be limited to 300 HV5 and this could be predicted through simulation work. Defect-free joining of 12 mm-thick X2CrNi12 stainless steel plates is also possible by laser welding without filler metal but in this case fused metal microstructures are obtained with a hardness of about 300 HV5. 4. The major shortcoming of this stainless steel is the tendency for grain coarsening at the heat-affected zone close to the fusion line if the heat input during welding is not properly controlled. Grain coarsening has no adverse influence on tensile properties or bend properties, but the heat-affected zone impact toughness for sub-zero temperatures may be disappointing and this certainly depends on the amount of grain coarsened microstructures. One technique for improvement consists in selecting a wide angle for the plate preparation leading to a weld geometry with strongly inclined fusion lines with regard to the plate surface. Indeed experience gained outside the present research project learns that in this case notch impact toughness test samples systematically yield improved data. As this increases the width of the weld it may somewhat adversely affect the atmospheric behaviour of welds provided with a coating. In general though, an appropriate coating should protect the weld for an estimated period of several decades. 5. However, various welds have proven that adequate heat-affected zone impact properties are achievable down to –40 °C or even lower and this is an encouraging result. Despite the complexity of microstructures sampled by notches located at the heat-affected

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zone, microscopic investigations have shown indeed for these ‘high quality’ welds that grain coarsening could be restricted to microstructures with an ASTM grain size number of 5 or higher. The fact that a correlation between good toughness and applied cooling rates could not be defined, which would have helped in indicating ‘safe’ heat input ranges for various plate thicknesses, was mostly disappointing. In addition, research through weld simulation showed that toughness at the high temperature heat-affected zone was independent of the selected cooling rates. Differences between heat-affected zone toughness of various real welds were therefore believed to be rather attributed to other aspects that are uncontrollable during welding, see next conclusion. 6. The observation of scatter in heat-affected zone impact toughness and of disappointing fusion line CTOD fracture toughness is another point of concern. It is strongly believed that this can be accredited mainly to the heat treatment of the base metal that allows only slight deviations, but partly also to the presence of small added quantities of boron. If this problem is solved, then most of the actual welds should reveal sufficient toughness allowing their use at temperatures as low as –20 °C or even –40 °C. 7. The variation in base metal heat treatment during production can also be the cause of scatter in tensile properties. A tighter tolerance on tensile strength would limit the risk of weld metal mechanical undermatching in case of austenitic filler metal. This can be achieved by limiting the actual base metal strength to 600 MPa. 8. Fatigue properties and resistance against atmospheric attack of ‘clean’ X2CrNi12 ferritic stainless steel welds are attractive, provided weld defects are omitted and, for the former, all excess of weld metal is removed appropriately or the geometric transition from weld to base metal is suitably smoothened. Its interesting position with regard to non alloy structural steels and expensive austenitic, martensitic and duplex stainless steels has thus been proven by the project described in this document as it combines productivity with rather low investments and maintenance costs resulting in long-term attractive solutions.

ACKNOWLEDGEMENTS The authors acknowledge the assistance of all colleagues at the Research Centre of the Belgian Welding Institute. Also the support of IWT and of all members, i.e. Industeel, Aelterman, Bombardier Eurorail, Buyk Steel Constructions, COEK Engineering, CMI Energy Services, ESAB, Infrabel, Lincoln Smitweld, Ministerie van de Vlaamse Gemeenschap (Metaalstructuren), OCAS, Remytole, TUC-Rail, University of Gent, VCLS/ CPS, VITO, Vyncke and WTCM is gratefully acknowledged. Without their help and technical support the present research could never have been accomplished so successfully. Welding in the World, Vol. 53, n° 9/10, 2009 – Peer-reviewed Section

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