duplex stainless steel for bridges construction: comparison between saw and laser-gma hybrid welding

DUPLEX STAINLESS STEEL FOR BRIDGES CONSTRUCTION: COMPARISON BETWEEN SAW AND LASER-GMA HYBRID WELDING R123

Welding in the World, Vol. 54, n° 5/6, 2010 – Peer-reviewed Section

1 INTRODUCTION

In recent years, austenitic – ferritic stainless steels (duplex stainless steels – DSS) have been extensi-vely employed in both bridges and footbridges for deck construction [1]. Major key point of that mate-rial, compared with the high strength C-Mn steels (e.g. S460) traditionally employed in such structures, is its resistance against corrosion, as high to allow the use

even in marine/industrial environment without surface protection and with very low maintenance cost during its life cycle.

The favourable combination of mechanical properties, manufacturability and durability makes competitive DSSs compared to the analogue welded construction made of C-Mn steel, if all factors that infl uence the structure life cycle cost are taken into the evaluation, beyond the merely mechanical behaviour of the mate-rial [2]. In fact, for structures made of DSSs, over time, the convenience relies on the lower costs for surveil-lance and maintenance and the absence of the initial anticorrosion treatment [3]. Obviously, the best eco-nomy is achieved if the structure is to be in service in severe aggressive environment (e.g. marine).

Up to now, the use of DSS in bridges was confi ned to pedestrian and/or small span structures. In sight of the possible employment in more challenging projects

DUPLEX STAINLESS STEEL FOR BRIDGESCONSTRUCTION: COMPARISON BETWEEN SAW

AND LASER-GMA HYBRID WELDING

M. Fersini S. Sorrentino G. Zilli

ABSTRACT

Using duplex stainless steels for bridge decks would be a major step forward in providing durable, low maintenance structures, exploiting both their corrosion resistance and high mechanical properties, while fulfi lling the required structural safety performances. In this work we investigated the possibility to replace the longitudinal welded joint between the deck plate and the girder trapezoidal stiffeners, performed by SAW, with a comparable one fabricated by hybrid laser welding (LB-GMAW). Potential advantages that could be achieved are, among others, productivity improvement, better detail fatigue resistance and lower distortion of the welded structure. The main differences obtained using SAW and LB-GMAW for the specifi c structural detail made in duplex stainless steel (UNS S32205) are presented. Technological, metallurgical features and weld integrity were mainly investigated. The utilization of the LB-GMAW technique resulted in welded joint fully penetrated, differently from the case of SAW, a higher pro-ductivity and reduced distortions. From a metallurgical point of view (i.e. austenite-ferrite phase balancing), results are still subjected to be improved but acceptable fi gures were already achieved.

IIW-Thesaurus keywords: Combined processes; Duplex stainless steels; Fatigue strength; Laser welding;Mechanical properties.

R123


Dr. Maurizio FERSINI ([email protected]), Head of department, Dr. Stefano SORRENTINO ([email protected]) and Dr. Giuliana ZILLI ([email protected]) are all with Centro Sviluppo Materiali S.p.A, Roma (Italy).

Doc. IIW-1995-09 (ex-doc. IV-960r1-08), recommended for publication by Commission IV “Power Beam Proc-esses.”

DUPLEX STAINLESS STEEL FOR BRIDGES CONSTRUCTION: COMPARISON BETWEEN SAW AND LASER-GMA HYBRID WELDINGR124


2 MATERIALS AND WELDABILITY

DSSs are ferrous alloys with two-phase microstructure; austenite and delta ferrite roughly in the same propor-tion. They combine outstanding technological proper-ties, like higher strength compared with austenitic stain-less steels, better resistance to the uniform corrosion, intergranular corrosion, pitting corrosion, outstanding resistance to stress-corrosion cracking and adequate weldability. We address to the large existing literature on DSSs for specifi c topics. We strongly recommend, particularly, reading of reference [5], which includes a large bibliography.

The candidate DSS grade for the bridge deck construction is a commercial 2205 gradation, with composition close to the upper boundary (Uranus 45N Mo®-2306) of the win-dow allowed by the relevant standard, compared to the median chemistry. This grade, nevertheless, lies in either the standard UNS S32205 or EN 1.4462 designations. The PREN index (PREN = Cr% + 3.3 Mo% + 16 N%) value is the same or better than 34, higher molybdenum and nitrogen content than standard 2205 are included for the best corrosion resistance even for very thick plates. 2205 grade is the DSS more employed, for which a large database of properties has been developed over years.

The actual chemical compositions are reported in Table 1, for the plates and the stiffeners respec-tively, while in Figure 1 the base material (BM) typical microstructure is shown.

The actual microstructure had slightly more auste-nite (about 45 % ferrite) than the normal balancing 50 %-50 %. Nitrogen addition ensures a better weld-

(e.g. Messina Bridge), it has been necessary to improve the knowledge of the behaviour of DSS large welded structures, in both the environmental and load sce-narios typical of the target service. Only in this way it could be possible to assess costs saving and to inte-grate the existing structural design codes for bridges (Eurocode – EN 1993-2:2006) with data applicable to austenitic-ferritic steels (today included only in the general structural code EN 1993-1-4:2006). These cru-cial objectives were pursued by a European research project funded by the Research Fund for Coal and Steel (RFS-CR-04040 “Bridgeplex”).

Of course, welding technology plays a crucial role in the structure manufacturing. Joint soundness, cost saving and welding process robustness and reli-ability are always of extreme importance. For very large structures, prefabrication of sections in shop is widely employed, where the goal is to maximize the use of mechanised process with high productivity for long lon-gitudinal joints. For that purpose, the Submerged Arc welding process (SAW) is widely used.

In the present work, beyond the Bridgeplex planned acti-vities, the Laser-GMA (LB-GMAW) hybrid welding process has been investigated as candidate for replacing the SA welding referring to the longitudinal joints between trape-zoidal stiffeners and a large span bridge deck.

Due to the intrinsic behaviour of the LB-GMAW process, several potential advantages could be achieved:– the capability to obtain a fully penetrated joint (for a specifi c structural detail) by welding from one side only (whilst by SA technique the same detail has quite always partial penetration);– better welded detail fatigue strength and therefore the opportunity for lightening the structure and saving expensive material;– elimination of potential starting locations for crevice-corrosion;– narrower welded area, lower residual stresses and therefore less risk of stress corrosion;– absence of thermal distortions (while signifi cant in SAW joints and diffi cult to recuperate on DSSs by ther-mal methods), that means better precision and con-struction quality for the structure;– increased productivity, reduction of expensive con-sumable material.

The present work is focused on the differences be-tween SA and LB-GMA welds for the same structural detail. Technological, metallurgical issues and weld soundness were mainly investigated. SAW data were obtained from Bridgeplex research project, available documentation [4].

Table 1 – Nominal and actual chemical composition for Uranus 45N Mo steel (UNS S32205 – 2306)

Grade C S P Si Mn Cr Ni Mo N PREN

2306 (7 mm)a 0.025 0.0006 0.024 0.53 1.69 22.49 6.15 3.03 0.191 35.5

2306 (12 mm)a 0.023 0.0004 0.025 0.50 1.63 22.50 6.01 3.07 0.187 35.5

EN 1.4462 < 0.030 < 0.015 < 0.035 < 1.0 < 2.0 21.0 ÷23.0 4.5 ÷ 6.5 2.5 ÷ 3.5 0.10 ÷ 0.22 -

UNS S32205 < 0.030 < 0.02 < 0.03 < 1.0 < 2.0 22.0 ÷23.0 4.5 ÷ 6.5 3.0 ÷ 3.5 0.14 ÷ 0.20 -a Mass %, ladle analysis, Fe % balancing.

Figure 1 – Austenitic-ferritic Uranus 45N Mo steel microstructure (near surface, dark: ferrite; bright:

austenite; courtesy of Arcelor Industeel)



ables are reported. SAW data come from the Bridge-plex project documentation.

3 WELDED JOINT CONCEPTION –FATIGUE RESISTANCE

Usually, large span bridge structures employ the ortho-tropic deck concept, in which the deck plates are stiffen-ed by girder longitudinal ribs with trapezoidal geome-try. Further, there are, at regular distance, transversal beams welded to both the deck and the trapezoidal stiffeners; to allow the beams to be connected to the profi les, large holes are cut out in shop for the profi les intersection (Figure 2).

ability due to the improved stability of the phases in HAZ. For the structural detail under investigation, pla-tes of 7 mm (stiffener ribs) and 12 mm (deck plates) in thicknesses were provided in the solution annealed condition.

In Table 2 the nominal mechanical properties are re-ported. Mechanical performances are comparable to a 460 MPa class C-Mn steel.

For arc welding of DSSs, consumables slightly more alloyed in austenite former elements (2209 grade) or highly alloyed (superduplex grades) are recommen-ded. Superduplex grades (e.g. 2510) allow for welded joints with better pitting corrosion resistance. Both consumable types were included in LB-GMAW expe-riments. In Table 3 the nominal properties of consum-

Table 2 – Uranus 45N Mo steel nominal mechanical properties (UNS S32205 – 2306)

Grade Properties Rp0.2 Rm A

(MPa) (MPa) (%)

2306 Minimum a > 460 > 680 -

EN 1.4462 (12 ÷ 75 mm) standard > 460 > 660 > 25

UNS S32205 (> 6 mm) standard > 450 > 655 -a Source: steelmaker.

Figure 2 – Orthotropic bridge deck concept (in detail)

Table 3 – Properties of the welding consumables

Consumables Process C Si Mn Cr Ni Mo N PREN FN

EN 12072S22 9 3 N L (wire)

SAW 0.03 0.5 1.4 22 9 3 0.15 34 30-50EN 760

SA AF 2 DC (fl ux)

AWS 5.9-93 ER2209 LB-GMAW < 0.020 0.5 1.6 23 9 3.2 0.16 40 (ferrite %)

Zeron100X™ LB-GMAW 0.015 0.4 0.7 25 9.8 3.7 0.22 ≥ 40 a

Mass %; Fe balancing. a Cu 0.6 %, W 0.7 %.



Focusing on a large span bridge, the structural design carried out the nominal dimensions for the investigatedstructural detail, subjected to the loading scenariotypical for the application.

– deck plate: 12 mm;– girder trapezoidal rib: 7 mm.

As the trapezoidal profi le was quite limited in dimension (envelope 300 x 300 mm), it was evident that no weld-ing activity might be performed from the internal side to seal the root area. Moreover, full penetration could not be ensured by welding from the external side only without root backing (either permanent or temporary). Backing application appeared to be quite inapplicable, as well.

Looking at the loading path, it is worth to note that the stiffening ribs themselves are not subjected directly to signifi cant loads. To withstand the prevalent longitu-dinal static load, it is not required that the joints must have full penetration. It is possible to fulfi l the static resistance requirement even with incomplete penetra-tion (Figure 5), but, of course, the total throat thickness must be equal or above the trapezoidal rib thickness (e.g. 7 mm).

The orthotropic deck concept features strong benefi ts (low weight, good aerodynamic), but it is complicate to be manufactured, as made by a large number of complex welded details. Of course, the concept works for both DSS and traditional C-Mn steels. Concerning the in-service performances of such structure design, the presence of many complex welds, combined with variable loading due to the vehicles traffi c, make the prevention of fatigue damaging of the utmost impor-tance. In fact, for large suspended bridges, it is gene-rally recognised that the bridge design constraints comes from fatigue concerns.

The investigated structural detail (Figure 3) was the lon-gitudinal welded T-butt joint between the plate deck and the trapezoidal stiffening ribs.

Such details are usually welded in shop during the prefabrication stage, when deck sections are build-up before in-site erection. The preferred practice relies on two SA heads welding simultaneously in fl at position from the external side. SAW process may ensure the required productivity, weld integrity and the prescri-bed mechanical properties. The trapezoidal rib comes always from a cold formed plate, the bevelled edges being obtained by thermal cutting and machining. For the purpose of the present study, it was not investi-gated any particular bevel geometry: the weld prepara-tion was simply arranged by putting in contact the deck plate surface with the square edge obtained after rib cutting (Figure 4).

Figure 3 – Investigated structural detail (circled)

Figure 4 – Bevel preparation for both SAWand LB-GMAW (welding from the right side)

Figure 5 – Structural welded detail and joint groove penetration requirements (EN 1993-2:2006)



4 EXPERIMENTAL ACTIVITY

SA welded details were manufactured by OMBA Engi-neering S.p.A. within the Bridgeplex project, after ade-quate WPS development. The same structural details were welded by LB-GMA process with the facilities of Centro Sviluppo Materiali S.p.A. Welding Laboratory, in Rome.

The laser system consisted of a solid state Nd:YAG laser source, pumped by laser diodes, with a maximum power of 4.4 kW. The laser beam was transmitted to the workpiece by an optical fi bre of 600 μm in diameter, focused with a lens of 120 mm focal length after colli-mation through a lens with 200 mm of focal length. The welding head was fi tted with a MIG torch connected to an inverter generator rated 500 A, with full digital control. Laser spot and wire tip were kept quite at the same point. The welding head was manipulated by a 6-axis robot, the path pre-programmed by a numerical control. The protective gas was supplied by the MIG torch standard nozzle.

Welding trials were carried out on 500 mm long sam-ples, made of a 12 mm thick base plate coupled with a trapezoidal stiffener 7 mm thick. The stiffener edged was prepared by shearing, tacking on the base plate was made on the extremities. The welding process was performed in 1G/PA position.

Welding parameters were explored during the experi-ments searching the best combination for ensuring weld seam soundness, adequate productivity and favourable

Such choice is not optimal for the detail fatigue re-sistance, in which the dynamic loading path is quite particular (out-of-plane bending): loading fluctua-tions promoted from the vehicles traffi c induce struc-ture deformations that “open” the welded joint root (Figure 6).

At present, there are no fatigue design codes including structural details made of DSS (or stainless grades, as well). It is generally recognized that the use of design data pertinent to the same detail made of C-Mn steel is a safe approach (such practice is recommended in EN 1993-1-9:2005 standard).

For the described structural detail, a fatigue resistance class is set accepting a limit amount for lack of pene-tration (max.), while the other geometric fi gures (e.g. gap between members) are strictly controlled. In such a way, the design fatigue strength is 71 MPa at 2 mil-lion of cycles (if, beyond the incomplete penetration, other requirements are not fulfi lled, one must assume 50 MPa only instead of 71 MPa).

Such circumstance greatly infl uences the structure design, as the investigated detail is present everywhere in the deck, built with a very expensive steel. If a new welding technique could be employed robustly, with confi dence of full penetration, it may be possible to validate a higher fatigue resistance class for full pene-trated joints, with evident advantages. This, in principle, could be transferred to welded structures made of C-Mn steel, but with cheaper material the technical-economical convenience appears less certain.

Figure 6 – Out-of-plane bendingof the orthotropic deck



Table 4 – Optimized welding proceduresfor SAW and LB-GMAW (Zeron100X wire)

Steel UNS 32205 UNS 32205

Product type Plate Plate

Joint type T-butt T-butt

Welding position PA PA

Nominal thickness7 onto12 mm

7 onto12 mm

Backing - -

Cleaning practiceBrushing

or grindingBrushing

or grinding

Welding processLaser – GMA

weldingSAW

Pass 1 1

Laser Nd:YAG Rofi n Sinar DY044 -

Operating mode CW -

Optical fi bre diameter [μm] 600 -

Collimator [mm] 200 -

Focal length [mm] 120 -

Spot/wire single single

Laser actual power [W] 4 200 -Laser beam tilting from verti-cal direction [°]

75 -

Gas fl ow [l/min] 25 -

Consumables See Table 3 See Table 3I

Polarity DC+ DCEP

Interpass - -

Travel speed [m/min] 1.2 0.66

Focus position [mm] -3.0 -

Wire diameter [mm] 1.2 3.2

Gas blend Ar + 2 %CO2 -

Current [A] 390 400-410

Voltage [V] 29 26

Wire feed rate [m/min] 12 n.a.

Preheating [°C] 20 20

Gross heat input(ν = 1) [kJ/mm]

0.77 1.10

(SAW data courtesy of OMBA Engineering SpA)

In the application under investigation, it is unrealistic to perform a PWHT (solution annealing) to restore the right ferrite – austenite proportion.

LB-GMA welding, due to the synergetic effect of the electric arc thermal source, is generally recognized to have higher heat inputs than laser alone, and the possibility of wire addition and the bevel preparation more open should allow for less BM diluted into the FZ. These potential advantages could counter balance intrinsic weakness of laser applied on DSS. In the fol-lowing the main results are discussed.

Operating behaviour

The bevel preparation was identical for both SA and LB-GMA welding. Both welding methods produced

weld microstructure. The fi nal WPS was established after validation of the most promising parameter com-bination by visual testing of the welds, NdT and metal-lographic survey. Welded joints were subjected to the exams included in EN 15614 standard for weld quali-fi cation, like VT, PT, hardness testing and microstructu-ral investigations. The welded zone was investigated on macrosections by optical microscopy for evaluating the seam morphology. Microstructural analysis was carried out at 200x magnifi cation on polished (with 1 μm dia-mond) and etched samples. The proportion between austenite (γ) and ferrite (α) phase was assessed by both manual point counting and automatic image analysis. Local chemical analysis was performed in fused zone (FZ) and heat-affected zone (HAZ) by EDS (Energy Dis-persive X-ray Spectroscopy) at a Scanning Electron Microscope (SEM) facility.

5 RESULTS AND DISCUSSION

Fundamental process fi gures for the optimized WPSs are reported in Table 4.

Single pass laser (or electron beam) welding on DSS is generally recognized to be very diffi cult [5]. It is widely accepted that in both FZ and HAZ the metallurgical structure deviates considerably from the desirable phase proportion (α/γ ~ 1) after welding, while by arc welding it may be obtained with some effort. In nor-mal conditions, laser welds have austenite phase in FZ below 30 % (frequently as low as 15-20 %), which is the lower limit often prescribed in standards for an austenitic-ferritic microstructure capable of decent mechanical properties and corrosion resistance. Super-duplex grades (SDSS), due to higher content of auste-nite former elements may display 30-35 % of austenite in FZ [5] in the as-welded state. In fact, laser welding cannot employ easily the techniques developed for arc welding to withstand the tendency to the ferrite predo-mination:

– the heat input is always quite low, especially for low thicknesses, irrespective of procedure adjustments (Δt1200-800°C parameters may be as low as 1 s);

– BM dilution into the weld is full (Rd = 100 % – joint without fi ller addition) or very high; the benefi cial effect of over alloyed consumable is nearly lost and the che-mistry of the FZ is quite unchanged compared to BM;

– the generally positive infl uence of nitrogen added on the protective mix is uncertain, as it cannot dilute uni-formly in thick welds (the effectiveness at the root is low). Moreover, given a percentage amount in the gas mix (e.g. 5 %), porosity in FZ may arise easier than in arc welds due to the higher cooling rates and the nar-rower weld pool. However, BM includes a considerable amount of nitrogen, so, given the considerable dilution that occur, some nitrogen may be retained in FZ after solidifi cation;

– preheating should be very high (e.g. 250 °C) and therefore impractical.



Non-destructive testing

Both kinds of joints were free from signifi cant imperfec-tions after both visual and penetrant testing (VT, PT). Destructive testing for possible presence of volumetric

steady and repeatable processes, with very few spat-ters (dangerous on DSS) during the LB-GMAW pro-cess. Limited gaps (about 1.5 mm) between the plate and the stiffeners could be tolerated without parame-ter correction due to the use of the fi ller metal. SAW process, optimized for adequate penetration and cor-rect heat input (1.10 kJ/mm) for obtaining a balanced microstructure, retained good productivity (0.66 m/min). About this fi gure, LB-GMAW process was much more performing (1.2 m/min). As often found, the laser-based process displayed a quite narrow operational window, due to the various requirements that must be fulfi lled at one time (productivity, full penetration, reinforcement shape, appropriate heat input). The gross heat input for LB-GMAW was 0.77 kJ/mm (the net one being possibly much lower), then at the lower limit of the recommend-ed range for DSS welding. The chemistry infl uence of the wires investigated on the LB-GMA process para-meters was inappreciable.

Welded joints morphology

Both weld types had very regular appearance. SAW welds merged more smoothly to the parent plates. Throat thickness was 7 mm, lack of penetration about 2 mm, then fulfi lling with minimum margin FAT 71 fati-gue class requirements (Figure 7 and Figure 8).

LB-GMA welds with optimize procedure had complete and stable penetration at the root. Root area had few oxide scale, despite the absence of forming gas at the root (that could not be provided in a realistic appli-cation). This behaviour is attributable to the very little reinforcement and the rapid solidifi cation (Figure 9 and Figure 10). Such circumstance is not optimal for root corrosion resistance, but it must be noted that the root reinforcement metal is exposed to the more favourable internal environment of the closed box. (for SAW HAZ only, this occurs as well).

LB-GMA welds had reinforcement with convex shape at both cap and root side. As this was not the ideal morphology, it was not possible to improve the proce-dure for overcoming such issue without impairing pro-ductivity and/or weld microstructure. The FZ was, of course, narrow and elongated, with quite no distortion for the whole sample. HAZ was quite inappreciable at low magnifi cation (Figure 11).

Figure 7 – SA weld cap side appearance (UNS S32205, v = 0.66 m/min. Q = 1.10 kJ/mm,

after sand blasting)

Figure 8 – SA welded joint macro section (UNS S32205, v = 0.66 m/min, Q = 1.10 kJ/mm,

oxalic acid electrolytic etching)

Figure 9 – LB-GMA weld cap side appearance (UNS S32205, v = 1.20 m/min, Q = 0.77 kJ/mm,

ER2209, after seam pickling)

Figure 10 – LB-GMA weld root side appearance (UNS S32205, v = 1.20 m/min, Q = 0.77 kJ/mm,

after seam pickling)



relevant for the specifi c zone being investigated. In Table 5 are reported the results for the two joint types considering all consumables.

LB-GMA welded FZs were analysed at both cap and root location, while SA welds were analysed at a single location (mid FZ). Some elements (like nitrogen) cannot be detected with good precision by EDS technique, but that was the only method suitable for very narrow FZs.

As predicted, the SA weld metal was correctly richer of γ-former elements and the WM PREN index was high-er than the BM, ensuring adequate resistance against pitting corrosion. In LB-GMA welds, the FZ was leaner in austenite formers (even considering the respective welding consumables), due to the higher BM dilution into the WM. It is worth to note that the joint root area had a chemistry more close to the fi ller wire one, while the cap WM was made essentially of BM. This is parti-cularly true for the ER2209 electrode, while Zeron100X (SDSS) wire allowed for more FZ alloying and higher PREN index fi gures.

Welds microstructure

A balanced microstructure is the crucial feature for obtaining DSS welds with both mechanical and tech-nological properties fulfi lling the relevant application standards.

In Table 6 is reported the phase proportion measured after fi nal WPS development.

At the optical microscope, austenite and ferrite are the only phases observed. There is no evidence of brittle intermetallics (σ, χ etc.), nor of carbide. Images from Figures 12 to 17 make possible to have an idea at fi rst glance of the predominance of either ferrite or austenite and the absence of further constituents.

imperfections was performed by examination of several metallographic sections for each type. SAW joints were free from any internal imperfections, while in LB-GMA joints some small isolated pores with small diameter (0.1–0.4 mm), were revealed.

WM chemical analysis

FZ chemical analysis was important for estimating phase balancing by constitutional diagrams and the welded joint corrosion resistance by the PREN index

Figure 11 – LB-GMA welded joint macro section (UNS S32205, v = 1.20 m/min, Q = 0.77 kJ/mm,

ER2209, oxalic acid electrolytic etching)

Table 5 – WM chemistry for the examined welded joints

Joint Method Si Mn Cr Ni Mo

SAW ZF EDS 0.57 ± 0.09 1.52 ± 0.19 23.22 ± 0.24 7.19 ± 0.24 3.45 ± 0.26

SAW a ZF Chemical n.a. n.a. 23.8 n.a. 3.1

LB-GMAW (ER2209) FZ (cap) EDS 0.56 ± 0.09 1.86 ± 0.19 22.72 ± 0.25 6.63 ± 0.25 3.44 ± 0.27

LB-GMAW (ER2209) FZ (root) EDS 0.54 ± 0.09 1.62 ± 0.19 22.56 ± 0.25 7.41 ± 0.25 3.42 ± 0.27

LB-GMAW (Zeron100X) FZ (cap) b EDS 0.55 ±0.09 1.56 ± 0.19 23.99 ± 0.26 6.55 ± 0.25 3.68 ± 0.27

LB-GMAW (Zeron100X) FZ (root) c EDS 0.47 ± 0.10 1.36 ± 0.19 24.30 ± 0.27 7.16 ± 0.26 3.59 ± 0.28

UNS S32205 - 2306 (7 and 12 mm) Chemical 0.53 ÷ 0.50 1.69 ÷ 1.63 22.49 ÷ 22.50 6.15 ÷ 6.01 3.03 ÷ 3.07

Mass %. b Cu = 0.52±0.20 %.a SAW N% = 0.18; “Bridgeplex” project. c Cu = 0.36±0.21 %.

Table 6 – WM Austenite-Ferrite balancing for SA and LB-GMA welded joints

Joint Method Standard Location α γ

SAWPoint

counting a

ASTM E562 (30 fi elds/25 points)

FZ 47.7 ± 3.3 % balancing

LB-GMAW (ER2209) AIA b Internal CSM FZ 72.8 ± 4.6 % balancing

LB-GMAW (Zeron100X) AIA b Internal CSM FZ 70.6 ± 4.1 % balancinga From “Bridgeplex” project.b Automatic image analysis.



Figure 12 – SA weld FZ microstructure (UNS S32205, v = 0.66 m/min,

Q = 1.10 kJ/mm, dark: austenite; bright: ferrite; oxalic acid electrolytic etching)

Figure 13 – SA weld microstructure near fusion boundary (left BM, right FZ)

(UNS S32205, v = 0.66 m/min, Q = 1.10 kJ/mm, oxalic acid electrolytic etching)

Figure 14 – LB-GMA weld FZ microstructure with ER2209 wire (UNS S32205, v = 1.20 m/min, Q = 0.77 kJ/mm, dark: austenite; bright: ferrite;

oxalic acid electrolytic etching)

Figure 15 – LB-GMA weld HAZ microstructure with ER2209 wire (bottom BM, top FZ) (UNS S32205,

v = 1.20 m/min, Q = 0.77 kJ/mm, oxalic acid electrolytic etching)

Figure 16 – LB-GMA weld FZ microstructure with Zeron100X wire (UNS S32205, v = 1.20 m/min,

Q = 0.77 kJ/mm, dark: austenite; bright: ferrite; oxalic acid electrolytic etching)

Figure 17 – LB-GMA weld HAZ microstructure (bottom BM, top FZ) with Zeron100X wire (UNS S32205, v = 1.20 m/min, Q = 0.77 kJ/mm, oxalic

acid electrolytic etching)



performed similarly (considered with the measurement uncertainty). It cannot be excluded that marginal WPS enhancement could improve the austenite content until the 35 % limit (e.g. minimum requirements in DNV-OS-F101 offshore standard [6]). It is interesting to note that LB-GMA welds had more dispersed values be-tween the various WM locations (cap, root), as shown by AIA, probably because of both not uniform dilution and sharp thermal fi eld, like instead happened in SAW joints. Looking at the reported chemistries, it was evi-dent the negative infl uence of the low heat input for a FZ correctly alloyed with Ni (at least at the root), and the loss of nitrogen as well, which cannot be estimated easily.

From the reported images, it can be noted that the austenitic phase was very fi ne in FZ of LB-GMA joints, and the HAZ extension could be appreciated at very high magnifi cation only.

Vickers hardness surveys (profi les) were carried out on both joint types welded zones with 10 kgf load, mee-ting EN 15614-1 standard. In Table 7 are reported the relevant results.

LB-GMA weld had slightly higher values at FZ, particu-larly if welded with Zeron100X wire. In any case, peak values were well below the limits prescribed in most recognized standards (e.g. max. 350 HV in DNV-OS-F-101), so a good weld ductility could be predicted.

It was evident from the gained results that the LB-GMA welding process could prevent most of problem asso-ciated with the application of the laser technique, which its intrinsic weaknesses when applied on such materi-als. Optimized welded joints had interesting properties such as better phase balance, higher productivity, fully penetrated joint, particularly when compared to DSS welds made by laser alone.

Productivity and weld integrity may be favourably infl u-enced by the LB-GMAW technique, while mechanical properties (that at this stage could be inferred from microstructural analysis only) and corrosion resistance may be of interest. Higher austenite phase content (at least 35 %) could give better reliability to the process. Corrosion credits for using a superduplex grade con-sumable must be assessed by specifi c testing.

The fatigue properties will be assessed by appropriate fatigue tests on the welded details, also on full scale stiffened panels. This is the purpose of the future work, for which several prototype stiffened panels were manufactured, to complete the test series already per-

Inside LB-GMAW joints, austenite (even if very fi ne) was present at primary ferrite grain boundary in roughly the same amount as of SA welds, while the interior was poorer of gamma phase.

Phase proportion in HAZ was unbalanced towards more ferrite in all joint types. Here differences between SAW and laser based joints were less steep, also due to the laser HAZ narrowness and the benefi cial nitrogen effect. However, deviation from the optimal proportion was not important. No grain growth was observed.

As reported by some authors [7] in works on either laser or electron beam welding of DSS, chromium nitri-des were found in HAZ of LB-GMA welds, promoted by the very fast cooling and the nitrogen oversaturation of ferrite. The presence was not continuous along the whole HAZ, neither the amount so detrimental like in laser alone processing. It was attributable to the inter-mediate cooling rate of LB-GMA process (Figure 18).

The correct choice of SAW wire, fl ux and the appropri-ate WPS determined the ideal ferrite – austenite pro-portion.

As could be predicted from the heat input and the che-mical composition of FZ for LB-GMA welds, the ferrite phase was predominant, but with the Zeron100X wire we could approach the limit proportion 70 %-30 % that should ensure adequate mechanical properties and stress-corrosion resistance, while the resistance to pitting corrosion was not questionable. ER2209 wire

Figure 18 – LB-GMA Weld HAZ microstructure (bottom FZ, top BM) with Zeron100X wire

(UNS S32205, v = 1.20 m/min, Q = 0.77 kJ/mm, dark: ferrite; bright: austenite)

Table 7 – Vickers hardness survey (HV10) on SA and LB-GMA welded joints (UNS S32205 steel)

Joint Location VHN (average) VHN (range) N° of indentation

SAW FZ 260 257 ÷ 266 3

SAW HAZ 259 248 ÷ 264 6

LB-GMAW (ER2209) FZ (cap) 270 267 ÷ 274 3

LB-GMAW (ER2209) HAZ (cap) 259 254 ÷ 264 2

LB-GMAW (Zeron100X) FZ (cap) 283 274 ÷ 289 3

LB-GMAW (Zeron100X) HAZ (cap) 269 266 ÷ 272 2

UNS S32205 (7 and 12 mm) BM 250 247 ÷ 254 6

Note the chromium nitrides darker point-like band.



From a metallurgical point of view the results are still not optimal, but the target (30–35 % of austenite in FZ) appears to be quite close. It is probable that only the modifi cation of the bevel preparation could promote less dilution of BM into the weld metal and therefore a more favourable austenite-ferrite balancing. For LB-GMAW joints, there was evidence that the FZ had not uniform dilution comparing weld cap and root. Such a topic should be more thoroughly investigated insofar as it infl uences the prescribed corrosion resistance.

Although some limitations still present, LB-GMA weld-ing process, applied on DSS, appears to be enough robust and reliable, and able to achieve the potential advantages. The effectiveness in exploiting fatigue mechanical properties must be assessed by a specifi c fatigue testing programme, also performed on full scale details.

ACKNOWLEDGEMENTS

The authors wish to thanks the Bridgeplex consortium for the supplied data on arc welds.

REFERENCES

[1] Buzzichelli G., Maiorana E., Scasso M.: Gli acciai inossidabili austeno-ferritici: considerazioni sull’impiego nella costruzione di impalcati da ponte, Considerations on the use of duplex stainless steel in the construction of bridge decks, Rivista Italiana della Saldatura, 2004, no. 4, pp. 459-470 (in Italian).

[2] Zilli G., Fattorini F., Maiorana E.: Application of duplex stainless steel for welded bridge construction in aggres-sive environment, Duplex 2007 International Conference & Expo, Grado, 2007.

[3] Fanica A., Maiorana E.: UNS S32205 for bridge con-struction: an experience of application, Duplex 2007 Inter-national Conference & Expo, Grado, 2007.

[4] Fattorini F., Zilli G., Miazzon A., Maiorana E., Peultier J., Fanica A., Stangenberg H., Hechler O., Maquoi R.: Application of duplex stainless steel for welded bridge construction in aggressive environment (BRIDGEPLEX), RFCS Mid-term Report, 2006.

[5] Bonollo F., Tiziani A., Ferro P.: Evoluzione microstrut-turale di acciai duplex e superduplex in relazione ai pro-cessi di saldatura, Microstructural evolution of duplex and superduplex steels in relation to welding processes, La Metallurgia Italiana, 2005, vol. 97, no. 2, pp. 27-38, (in Italian).

[6] Offshore standard DNV-OS-F101: Submarine pipeline systems, 2003.

[7] Bonollo F., Tiziani A., Zambon A., Penasa M.: Laser beam welding of superduplex stainless steels, Proceedings of Conference Duplex Stainless Steels ’94, Glasgow, 1994, paper no. 108, The Welding Institute, Abington (GB).

formed on SA welded specimens within the Bridgeplex Project (Figure 19).

5 CONCLUSIONS

The study evidenced that the manufacture of complex carpentry welded structures, made of UNS S32205 DSS, may be approached successfully by using traditional arc welding methods (SAW in particular). It is possible, in this way, to achieve good productivity and mecha-nical-technological properties for the welded joints, meeting relevant specifi cations.

From the gained experience, it was evident that the Laser-GMA process may perform successfully in weld-ing a specifi c structural detail with reliable integrity.The synergetic effect of the combined arc and laser sources made possible to obtain with reliability a full penetrated joint with higher productivity (travel speed 1.2 m/min) and less concerns for eventual imperfec-tions on joint preparation. The reported results were achieved with realistic LB-GMA WPSs, in principle employable in shop for the prefabrication of similar structures made of DSSs.

Figure 19 – Welded stiffened panel (SAW)made of UNS S32205 steel being tested

for fatigue resistance

duplex stainless steel for bridges construction: comparison between saw and laser-gma hybrid welding

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