Introduction Offshore development has acceler-ated in recent years owing to the factthat more than 50% of undevelopedpetroleum deposits are located underthe ocean. In the offshore industryand in underwater oil and gaspipelines, underwater welding is al-ready a routine activity (Refs. 1, 2).The demand for underwater weldingprocesses that can produce quality wetwelds at greater depths, and on a vari-ety of materials, will continue to in-crease (Ref. 3). Underwater welding techniques canbe classified as follows: wet welding,dry welding, and local cavity welding.Wet welding occurs directly in aqueous

environments with no mechanical bar-rier between water and welding arc. Itwas established that significant costsavings and simplicity of the processmakes it possible to weld even themost geometrically complex struc-tures; therefore, underwater wet weld-ing is of increasing importance (Refs.4, 5). The most commonly used wetwelding techniques are shielded metalarc welding (SMAW) (Refs. 6, 7) andflux cored arc welding (FCAW). It wasacknowledged that wet flux cored arcwelding is promising in the future be-cause of much higher production effi-ciency and applying in the automaticwelding process (Refs. 8, 9). In order to meet the requirementsfor offshore structures, high-strength

steel (yield strength over 350 MPa) isrequired. The strength of the steel usedfor offshore structures is a very impor-tant factor (Ref. 10). Unfortunately,high-strength low-alloy (HSLA) steelsusually have carbon equivalents greaterthan 0.4% and show poorer weldability.At the same time, an aqueous environ-ment produces a lot of disadvantageouseffects (Ref. 11), such as the cooling ef-fect of the surrounding water, loss of al-loying elements, and considerableamounts of diffusible hydrogen (Ref.12). The cooling rate in wet welding ismuch higher than in dry welding, suchas in the temperature range from 800°to 500°C, it can rise sharply from 56° to415°C/s (Ref. 4). This causes brittleweld microstructures and high amountsof hydrogen porosity, which can becauses of crack formation. Susceptibilityto cold cracking is the main problem inwelding of HSLA steels and fabricationof dissimilar joints. Many researchers have attemptedto use special methods to avoid theseadverse effects. Many studies utilizedthe temper bead technique (Refs.13–15). A full welding procedure quali-fication without cracking has beencompleted for a base plate having acarbon equivalent of 0.44. However,this method is only suitable for repairof underwater structures, which limitsits application. In addition, insulatingmaterials (Refs. 16, 17) were used tocontrol cooling rates in underwaterwet welds. The research, taking intoaccount the insulating material, devel-oped an empirical relationship to pre-dict the optimized cooling rates and

WELDING RESEARCH

Preliminary Investigation on Real­Time InductionHeating­Assisted Underwater Wet Welding

A unique process that combines induction heating and flux cored arc wet welding to reduce cooling rates in real time was studied

BY H. T. ZHANG, X. Y. DAI, J. C. FENG, AND L. L. HU

ABSTRACT A novel real­time induction heating­assisted underwater wet welding process wasinvestigated. The addition of induction heating could reduce the cooling rate of thejoint in underwater wet welding. The macro­ and microstructures, mechanical prop­erties such as tensile, impact, and bending properties, and Y­slit restraint testingwere studied. The results showed the content of martensite (M) and upper bainite(BU) phases decreased, while the proeutectoid ferrite (PF) and acicular ferrite (AF)phases increased as the induction heating voltage increased. Mechanical propertiesof the joint were improved through addition of induction heating and fracture mor­phology with characteristic uniform dimples belonging to ductile fracture. The crack­ing ratio of Y­slit restraint testing was also decreased. Therefore, the susceptibility tocold cracking of the wet welding joint was improved.

KEYWORDS • Underwater Wet Welding • Induction Heating • Microstructure • Property

H. T. ZHANG (hitzht@163.com), X. Y. DAI (hitdxy29@gmail.com), J. C. FENG, and L. L. Hu are with the State Key Laboratory of AdvancedWelding and Joining, Harbin Institute of Technology, Harbin, China; Shandong Provincial Key Laboratory of Special Welding Technology,Harbin Institute of Technology at Weihai, Weihai, China.

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times for underwater wet welds. Fox(Ref. 18) and Pope (Ref. 19) investigat-ed the water temperature and waterdepth influences on hydrogen-inducedcracking, microstructure, and mechan-ical properties in underwater wetwelding, and the importance of watertemperature and water depth, quench-ing, and diffusible hydrogen levels inunderwater wet welding have beendemonstrated. Postweld heat treat-ment (PWHT) is frequently used to re-duce hardened structure and allow hy-drogen to diffuse away from the weldmetal and heat-affected zone (HAZ)(Ref. 20). Szelagowski (Refs. 21, 22)used a H2-O2 cutting torch and an un-derwater high-velocity oxyfuel (UW-HVOF) thermal spraying device toserve as PWHT on wet welds. The hy-drogen content of the weld metal wasreduced and the bend testing resultshowed a higher plastic property.However, the control of heat inputcould not be accurate and efficient. In this paper, a novel real-time in-duction heating-assisted underwaterwet welding process was employed forthe first time. Induction heating couldreduce the cooling rates of the joint inunderwater wet welding, especially thet8/5 (the cooling time range from 800° to500°C) was extended. According towelding CCT diagrams, it reduced thehardened and brittle transformationproducts. That is, the content ofmartensite (M) and upper bainite (BU)phases decreased as the content ofproeutectoid ferrite (PF) and acicularferrite (AF) phases increased. Therefore,the purpose of this work was to develop

a novel method to obtain an excellentquality underwater wet welding joint.

Experimental Procedure Q460 steel (equivalent to Gr. 65 steelof AST-USA or E460DD steel of 630-ISO) delivered as rolled sections withthe dimensions of 300 × 90 × 8 mm wasused as the base metal. The single-Vweld groove had a 60-deg included anglewith a 2-mm root face and 1.5-mm rootopening. The chemical composition ofthe sheets is shown in Table 1. Prior towelding, the oxide layers on the surfacesof the plates were removed by stainlesssteel wire brushing and the weld zonewas degreased using acetone. The as-received plates were welded togetherwith the gas tungsten arc (GTA) andflux cored arc (FCA) welding processes.GTAW was used for the root pass to fixthe plates with 100-A DC and 20 V inair. Underwater wet FCAW was used forthe fill passes and optimized weldingparameters are listed in Table 2. TiO2-CaF2 type flux-cored wire with a diame-ter of 1.2 mm produced by Paton Weld-ing Institute was chosen. A schematic of the assembled de-vice is shown in Fig. 1. The devicecould be divided into two sections: un-derwater welding system and induc-tion heating system. The water in thetank was stationary and the waterdepth was 300 mm. A circular, 60-mm-diameter induction coil was installedbehind the welding gun in the weldingdirection and below the plates in thevertical direction. The welding gunand induction coil were fixed together

and moved at the same speed. The pa-rameter L — defined as the distancebetween the center of the coil and thewelding gun — was constant. The in-duction heating source had an outputvoltage of 70–550 V. Changing the in-duction heating voltage meant chang-ing the output power due to the con-stant system impedance. Type-K ther-mocouples with shielding were placedat different locations from the edge ofthe weld groove to measure the tem-perature profile. Four-channel dataloggers were used to record the tem-perature measurements with a sam-pling frequency of 25 Hz. The meas-urement method of the HAZ tempera-ture field was as follows: weld HAZwithout installed thermocouples wasfirst identified to be about 2.0 mmfrom the weld interface, then the ther-mocouples located at or near 2.0 mmfrom the weld interface were identi-fied as that representing the HAZthermal cycle (Ref. 23). A CCD camera with a frame rate of2000 frames/s was used to record im-ages of the arc behavior in order to in-vestigate the effect of the inductionmagnetic field. The metallographicspecimens of a typical cross sectionwere prepared vertical to the weld jointand all specimens were polished withSiC papers up to grit 1000, and ultra-sonically cleaned with acetone to re-move oil and other contaminants fromthe specimen surfaces. Etching with 5%nitric acid and alcohol solution for 3–4 swas used to reveal the weld beam. Themacro- and microstructure fracturemorphology were observed by opticalmicroscopy (OM) and scanning electronmicroscopy (SEM), respectively. Me-chanical property tests such as tensiletesting, impact testing, and bend test-ing were investigated to build an empiri-cal relationship between induction heat-ing voltages and mechanical properties.

Results and Discussion

Welding Process Stability

A welding arc is an electric dis-charge between two electrodes and aheated and ionized gas, called plasma(Ref. 24). Therefore, the arc stabilitycould be adversely affected as a resultof the magnetic field of inductionheating and eddy current. Figure 2shows the captured images of arc

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Fig. 1 — Schematic of the assembled device.

Table 1 — Chemical Composition of Q460 (not more than wt­%)

Base Metal C Si Mn V Ti Cr Ni Cu Mo

Q460 0.2 0.6 1.8 0.2 0.2 0.3 0.8 0.55 0.2

A BGun

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shape with parameter L and induc-tion heating voltages. Due to opti-mum parameters and flux-cored wire,the welding arc was steady during theunderwater welding process withoutinduction heating (Fig. 2A). While theinduction coil was installed, the arcstability was reduced. It was observedthat the parameter L played a majorrole in arc stability. When the param-eter L was 5 mm, the welding arc wasextremely unstable and even arc in-terruption appeared in Fig. 3A. At thesame time, when parameter L was in-creased to 20 mm, the welding arcshape was stable. Therefore, a contin-uous and uniform weld could be ob-served in Fig. 3B. Welding discontinu-ities, such as incomplete fusion andundercut, were not found. In addi-tion, the induction heating voltage af-fected arc stability and the arc stabili-ty decreased with increased voltage.To investigate the influence of volt-age on the joint, the parameter L wasfixed as 20 mm in the subsequentexperiments. Cross­Section Macrographs

Q460 sheets were underwater weld-

ed at a fixed welding parameter (Table2) and at various induction heatingvoltages ranging from 250 to 450 V inorder to clarify the effect of inductionvoltage on weld penetration. Cross-sectional macrographs of the jointswith different voltages are shown inFig. 4. According to the results, weldpenetration and HAZ increased withthe increasing voltage. As was known,the width of the HAZ depended pri-marily on heat input. The heat inputwas the sum of welding heat input andinduction heating. Therefore, the ef-fect of induction heating was equal toincreasing the welding heat input. Inaddition, the induction heating madethe temperature field of the weld zonerelatively more uniform.

Microstructure Characteristicsof the Joints

The HAZ for Q460 delivered asrolled sections mainly consisted of twodistinct zones: coarse-grained HAZ(CGHAZ) and fine-grained HAZ (FG-HAZ). Typical HAZ temperature vs.time profiles during the underwaterwet welding are shown in Fig. 5. Ac-

cording to the results, the cooling ratein wet welding was extremely higherthan in air welding. For instance, thecooling rate of the temperature rangefrom 800° to 500°C could rise sharplyto 100°C/s, which was more than thecritical cooling speed of martensiteformation. Figure 6 showed the optical mi-crostructure of the weld zone withvarious induction heating voltages.Based on the theory of welding met-allurgy, as the austenite phase wascooled down from high temperature,ferrite nucleated at the grain bound-ary at 770°–680° and then grew in-ward. This ferrite was proeutectoidferrite (PF), which is also called grainboundary ferrite (GBF). When thetemperature dropped to 500°, thetransformation of acicular ferrite(AF) occurred. The acicular ferritephase was a desirable phase becauseof the excellent plasticity and tough-ness characteristics (Ref. 25). As thecooling rate increased, the transfor-mation product changed to bainiteand martensite phase and reducedthe mechanical properties. The mi-crostructure of the weld zone in airwelding was composed of proeutec-toid ferrite and acicular ferrite phase.As was mentioned previously, the aci-cular ferrite phase had excellent plas-ticity and toughness, due to the inter-locking nature of the acicular ferriteand the fine granular size. Therefore,the mechanical properties were satis-factory. Compared to air welding, themicrostructure of underwater weld-ing was a mixture of lath martensite,upper bainite, and proeutectoid fer-rite — Fig. 6A. The bainite sheaf andmartensite lath nucleated and grewfrom prior austenite granular bound-aries. The formation of lath marten-site and upper bainite were detrimen-tal to the weld properties, owing tothe easy crack propagation paths. Asthe induction heating voltages in-creased (Fig. 6 B–D), the volume frac-tion of lath martensite and upper bai-nite decreased with ferrite phases in-creasing. Moreover, the lath marten-site and upper bainite phase disap-peared as the voltage reached 350 V.The transformation product (Fig. 6C)could change from upper bainite andlath martensite to acicular ferrite andproeutectoid ferrite. Therefore, themicrostructure of the weld metal was

WELDING RESEARCH

WELDING JOURNAL / JANUARY 2015, VOL. 9410-s

Fig. 2 — Captured images of arc shape: A — Without induction heating; B — L = 5 mm,250 V; C — L = 5 mm, 450 V; D — L = 20 mm, 250 V; E — L = 20 mm, 450 V.

Table 2 — Optimized Welding Parameters

Welding Voltage (V) Welding Current (A) Welding Speed (mm/min) Water Depth (mm)

26 160 145 300

A

B

C

D

E

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similar to air weld, the only differencewas the morphology of the proeutec-toid ferrite. Increased acicular ferritecontent in the microstructure im-proved cracking resistance, while up-per bainite and lath martensite dete-riorated the mechanical properties ofthe joint. The dimension of theproeutectoid ferrite was increasedwith the increase in the voltage. Asthe voltage was 450 V (Fig. 6D), themorphology of the proeutectoid fer-rite was coarsening and a ferrite sideplate (FSP) was found. Because of thelimited output voltage of the induc-tion heating system, the inductionheating was higher than 450 V, andthe microstructural evolution andmechanism are to be investigated inthe future. To understand the mechanism ofweld microstructural evolution, tem-perature vs. time profiles of differentinduction heating voltages without be-ing subjected to welding are shown inFig. 7. The parameters of t8/5, for a giv-

en hardenability steel, determined thehardenability of the transformationproducts, which should be taken intoconsideration to investigate the effecton susceptibility to hydrogen-inducedcracking. The data of temperature vs.time curves are shown in Table 3. Asthe induction heating voltage was 250V, the Tmax reached 412°C. Therefore,the microstructure of Fig. 6B was simi-lar to that shown in Fig. 6A becausethe t8/5 determined the transformationproducts. As the induction heatingvoltage was 350 V, the Tmax was in-creased to 609°C and the t8/5 was pro-longed at the same time. Therefore,the transformation products changedfrom upper bainite and lath marten-site to acicular ferrite and proeutec-toid ferrite due to the fact t8/5 was pro-longed. A comparison of temperaturecurves of 0 and 450 V is shown in Fig.8. It could be seen that the prolonga-tion of t8/5 was extremely obvious.That’s the reason for the evolution ofthe microstructure of the weld metal.

Figure 9 showed the optical mi-crostructure of the partially meltedzone with and without inductionheating. The red line was the weld in-terface of the joint. It could be seenthat lath martensite (M) and coarsen-ing Widmanstätten (W) structure waspredominant in the coarse-grainedHAZ in Fig. 9A. The ferrite phase pre-cipitated first in the coarse-grainedaustenite grain boundary, and thengrew into the austenite in the form ofreticular structure (also called Wstructure), resulting in splitting thematrix structure, even generating thecrack. And the lath martensite com-posed of vast coarse lath was benefi-cial for crack initiation and propaga-tion. Therefore, the mechanical prop-erties of the joint were decreased.However, as the voltage was 350 V,granular bainite was predominantand grain coarsening was relieved.Thus, the tendency to crack was de-creased, and the mechanical proper-ties of the joint were increased.

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Fig. 3 — Weld appearances: A — L = 5 mm, 250 V; B — L= 20 mm, 250 V.

Fig. 5 — Measured temperature vs. time curves of under­water wet welding without induction heating.

Fig. 4 — Cross­sectional macrograph of the joints: A — 0 V; B — 250 V; C— 350 V; D — 450 V.

Fig. 6 — The optical microstructure of the weld zone with various inductionheating voltages: A — 0 V; B — 250 V; C — 350 V; D — 450 V; E — in air.

A A

A

B

B

B

C

C

D

D E

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Mechanical Properties

Tensile Testing and FractureMorphology

Five prepared tensile specimensfrom each joint were performed usinga fully computerized tensile testingmachine with a loading rate of 1mm/min at room temperature to eval-uate the influence of various inductionheating voltages on the mechanicalproperties of the joint. The geometryof the tensile specimens and tensilestrength vs. voltage curves are shownin Fig. 10. The thickness of the speci-men was 4 mm. The tensile strength

of specimens without induction heat-ing was 444 MPa, about 82.2% of thebase metal (540 MPa). The tensileproperty increased gradually with in-creasing voltages. As the voltage was450 V, the tensile strength reached532 MPa, about 98.5%, and all thejoints fractured roughly in the HAZ.Joint efficiency increased from 82.2 to98.5%. In order to observe the fracturemechanism, SEM was carried out to an-alyze the fracture morphology. Figure11 shows the typical fracture surface ofspecimens with different voltages. Itcan be seen that a quasi-cleavage frac-ture mode was dominant in Fig. 11A,due to plenty of cleavage plane appear-

ance. The size of the cleavage plane wasrelated to the crack path. The largecleavage planes demonstrated very lowcrack propagation energy, while thesmall cleavage plane exhibited highercrack propagation energy (Ref. 26). Theformation of lath martensite and Wstructure in the HAZ was detrimental totensile property due to the easy crackpropagation paths. Once the crack oc-curred during the tensile test, it couldpropagate along the paths of lathmartensite and W structure rapidly.Therefore, tensile strength without in-duction heating was the lowest. Thedimple characteristics became predomi-nant as the voltage was increased in Fig.11B–D. Cleavage planes were in smallproportion while dimples were in largeproportion, as shown in Fig. 11B.Therefore, the fracture morphology hadthe characteristic of ductile fracture.While the voltage was above 350 V, thecleavage planes disappeared, instead ofuniform dimples, which was the typicalfeature of ductility. Nonmetallic inclu-

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Fig. 7 — Temperature vs. time profiles of different induc­tion heating voltages without being subjected to welding.

Fig. 8 — Compared temperature vs. time curves during un­derwater wet welding with various induction heating volt­ages.

Fig. 10 — Tensile strength vs. inductionheating voltage curves.Fig. 9 — The optical microstructure of the weld interface: A — 0 V; B — 350 V.

Table 3 — The Data of Temperature vs. Time Curves

Induction Heating Voltages (V) Tmax(°C) t8/5 (s)

250 412 0350 609 24450 712 35

A B

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sions phase were disorderly distributedin the inter-tear edges. Hence, tensileproperty of the joints improved toabout 98.9% of the base metal.

Impact Testing and FractureMorphology

Charpy V-notch impact tests wereconducted at 20°C on an instrumenteddrop weight impact tester. The speci-mens were extracted in the weld’s per-pendicular direction from the middlethickness of the as-welded specimenwith notches positioned at the centerof the weld metal. The impact energyof the joints was the average of fivespecimens. The geometry of theCharpy impact V-notch specimens andimpact energy vs. induction heatingvoltage curves are shown in Fig. 12.The thickness of the specimen was 10mm. The impact energy value of speci-mens without induction heating was36 J. As the voltage increased, the im-pact energy increased consistently. Fi-nally, as the voltage was 350 V, the im-pact energy could reach 68 J. The re-sults suggested that induction heatingcould increase the impact propertiesand the toughness of the joint. For the purpose of observing thefracture mechanism, SEM was used to

analyze the fracturemorphology. Figure 13shows the SEM micro-graphs of impact frac-ture surface morpholo-gies for different volt-ages. It can be seen thata cleavage fracture modeis dominant in Fig. 13A,due to a network ofcleavage steps known asa river pattern. Cleavagewas a low-energy frac-ture that propagated along well-defined low-index crystallographicplanes known as cleavage planes. Thebranches of the river pattern joinedin the direction of crack propagation.Meanwhile, the formation of lathmartensite and upper bainite was ad-verse to the toughness of the jointdue to the easy crack propagationpaths. Once the crack occurred, it rap-idly propagated in a straight linealong the lath martensite and upperbainite paths. Therefore, the impactenergy value without induction heat-ing was the lowest. However, the dim-ples started to appear and were pres-ent in small proportion when thevoltage was at 250 V in Fig. 13B.Feather markings, which are a fan-shaped array of very fine cleavagesteps on a large cleavage facet, are

present in large proportion. The apexof the fan points back to the fractureorigin. While the voltage was 350 V,the cleavage planes disappeared, in-stead of uniform dimples, which wasthe characteristic of ductile fracture.Some nonmetallic inclusions phaseshowed in a disorderly distributionand were surrounded by the inter-tear edges. The reason for these re-sults was that the acicular ferrite act-ed as the crack arrester and increasedthe crack propagation energy. There-fore, the impact property of the jointswas greatly improved from 36 to 68 J.

Bend Testing

Longitudinal three-point bend testswere conducted to measure the bendingductility at room temperature. The an-

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Fig. 12 — Impact energy vs. induction heating voltagecurves.

Fig. 11 — SEM images of the fractured surface of specimens withvarious voltages: A — 0 V; B — 250 V; C — 350 V; D — 450 V.

Fig. 13 — SEM images of the fractured surface of specimenswith various voltages: A — 0 V; B — 250 V; C — 350 V.

A

A

B

B

C

C

D

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gle of bending for the joints was the av-erage of three specimens. The geometryof the specimens and angle of bendingvs. induction heating voltage curves areshown in Fig. 14. The thickness of thespecimen was 5 mm. According to theresults, the angle of bending values ofspecimens without induction heatingwas 21 deg, which indicated ductilitywas very low. At the same time, the an-gle of bending values increased rapidlywith increasing voltages. Finally, as thevoltage was 450 V, the angle of bendingvalues reached 88 deg. The results sug-gested joint ductility had been increased.

Y­Slit Restraint Testing

Because of the high quenching ratecaused by the water environment andbecause large quantities of hydrogen arepresent, hydrogen cracking is one of themost severe problems in the underwa-ter welding of steel (Ref. 27). The crack-ing tests were carried out using a Y-slitrestraint test so that the intensity of therestraint could be related to actual fabri-cating conditions. The geometry ofspecimens subjected to Y-slit restrainttesting and the cracking ratio used as ameasure of the cracking susceptibilityare shown in Fig. 15. This is a ratio ofthe height from the root to the tip ofthe crack vs. the height from the root tothe surface of the weld metal.

The cracking ratio vs. induction heat-ing voltage curves are shown in Fig. 16.The carbon equivalent value of Q460steel was 0.6, which indicated the steelwas particularly sensitive to cracking,especially in underwater welding.

Therefore, the crack-ing ratio of speci-mens without induc-tion heating was

82%. However, as the voltage reached150 and 250 V, the cracking ratio de-creased rapidly to 45% and 22%, respec-tively. Finally, when the voltage wasabove 350 V, the cracking ratio reachedabout 10%. Typical weld cross sectionswith various induction heating voltagesare shown in Fig. 17. According to theresults, induction heating could reducethe cooling rate; therefore, the crackingsusceptibility was decreased.

Microhardness Profile

Vickers microhardness measurementacross the fusion zone was carried outwith a load of 100 g and load time of 10s. Results of hardness measurementsare shown in Fig. 18. The microhard-ness distribution indicated the mi-crostructural characteristics of the joint.Increased hardness values of the weldmetal confirmed these microstructuralchanges. The location of the HAZ wasdetermined by metallographic observa-tion and the hardness of the HAZ washigher than that of the weld metal. TheHAZ and weld metal hardness de-creased with increased induction heat-ing voltage. The maximum hardness ofHAZ without induction heating was 425HV, which was much harder than thatwith 250 and 450 V. The hardness val-ues of the weld zone with 250 and 450V were relatively uniform because thelath martensite and upper bainite co-tent decreased while the acicular ferriteand proeutectoid ferrite increased. Theresults indicated induction heating hada significant effect on the maximumhardness. The microhardness profile

across the weld indicated the mi-crostructural characteristics of the joint.Induction heating made the joint micro-hardness relatively more uniform.

Conclusion 1) A novel real-time induction heat-ing-assisted underwater wet weldingprocess was employed. The addition ofinduction heating could reduce the cool-ing rate of the joint in water environ-ment to improve the microstructuraland mechanical properties of the joint. 2) Arc stability was reduced with theaddition of induction heating. The pa-rameter L played a major role in arc sta-bility. As the parameter L increased to20 mm, the welding arc shape was sta-ble. A continuous and uniform weldjoint could be observed. 3) The content of martensite (M)and upper bainite (BU) phases de-creased while the proeutectoid ferrite(PF) and acicular ferrite (AF) phases in-creased as the induction heating voltageincreased. Mechanical properties, suchas tensile, impact, and bending proper-ties, increased as the induction heatingvoltages increased. 4) Cracking was examined via a Y-slitrestraint test. The addition of inductionheating could decrease the cracking ra-tio from 82 to 10%. Therefore, induc-tion heating could make cracking sus-ceptibility decrease.

1. Brown, R. T., and Masubuchi, K. 1975.Fundamental research on underwater weld-ing. Welding Journal 54(6): 178-s to 188-s. 2. Rowe, M., and Liu, S. 2001. Recentdevelopments in underwater wet welding.Science and Technology of Welding and Join-

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Fig. 14 — Angle of bending vs. induction heating voltagecurves.

Fig. 15 — Schematic of Y­slit self­restrained cracking test (mm).

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Fig. 17 — Typical underwater weld cross sections with various inductionheating voltages: A — 0 V; B — 150 V; C — 250 V; D — 350 V.

Fig. 16 — Cracking ratio vs. induction heating voltagecurves.

Fig. 18 — Hardness distribution along the joint.

A B

C D

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