laser beam welding of dissimilar ferritic/martensitic stainless steels in a butt joint configuration

Optics & Laser Technology 49 (2013) 125–136

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Optics & Laser Technology

0030-39

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n Corr

E-m

muhshi

journal homepage: www.elsevier.com/locate/optlastec

Laser beam welding of dissimilar ferritic/martensitic stainless steelsin a butt joint configuration

M.M.A. Khan n, L. Romoli, G. Dini

Department of Mechanical, Nuclear and Production Engineering, University of Pisa, Pisa 56126, Italy

a r t i c l e i n f o

Article history:

Received 13 August 2012

Received in revised form

10 November 2012

Accepted 17 December 2012

Keywords:

Laser-welding

Steels

Butt joint

92/$ - see front matter & 2012 Elsevier Ltd. A

x.doi.org/10.1016/j.optlastec.2012.12.025

esponding author. Tel.: þ39 502218137.

ail addresses: [email protected],

[email protected] (M.M.A. Khan).

a b s t r a c t

This paper investigates laser beam welding of dissimilar AISI430F and AISI440C stainless steels.

A combined welding and pre-and-postweld treatment technique was developed and used successfully

to avoid micro-crack formation. This paper also examined the effects of laser welding parameters and

line energy on weld bead geometry and tried to obtain an optimized laser-welded joint using a full

factorial design of experiment technique. The models developed were used to find optimal parameters

for the desired geometric criteria. All the bead characteristics varied positively as laser power increased

or welding speed decreased. Penetration size factor decreased rapidly due to keyhole formation for line

energy input in the range of 15–20 kJ/m. Laser power of 790–810 W and welding speed of 3.6–4.0

m/min were the optimal parameters providing an excellent welded component. Whatever the

optimization criteria, beam incident angle was around its limiting value of 151 to achieve optimal

geometrical features of the weld.

& 2012 Elsevier Ltd. All rights reserved.

1. Introduction

Laser welding with high power density, high degree of auto-mation and high production rate is extremely advantageous inindustrial applications as stated in [1]. Joints between dissimilarmetals are particularly common in components used in theautomotive, power generation, chemical, petrochemical, nuclearand electronics industries as described in [2]. The ability to usedifferent metals and alloys in a product provides the designer andproduction engineer with greater flexibility, and often results intechnical and economic advantages over components manufac-tured from a single material.

Welding of metals and alloys is an experienced subject,dissimilar welding represents a major scientific and technicalchallenge. Emerging new technologies increasingly require dis-similar metals and alloys to be joined. Most metals and alloys arevery weldable either in conduction or keyhole modes. Two factorsmost important to weldability are hardenability and susceptibil-ity of the hardened structure to cracking. Hardenability is relatedto the cooling rate of metals. The faster cooling rate tends toproduce higher hardness and hence, the hardened structurebecomes more susceptible to cracking. According to Kou [3], fora constant mass of the metal and the particular welding process

ll rights reserved.

and procedure, cooling rate depends only on weld preheattemperature as it can modify the energy input used to make theweld. It can also be changed to reduce the cooling rate. Fig. 1.1shows effects of combined welding and pre-and-post-weld treat-ment technique on mitigation of crack formation during welding.

Also, cracking may result from any of the following factor:restraint, weld shape, and material composition. Restraintis always present in any weld because as the weld solidifies itacquires strength but continues to cool and shrink. It is the degreeof restraint that becomes critical. Restraint relates to the welddesign, the weldment design, and welding procedure. Khan et al.[4,5] show that weld shape is also a function of the weld design,weldment design and welding procedure. Welding procedurerelates to the placement of welds or beads in the weld, the shapeof the beads and the shape of the finished surface as stated in [6].The third factor is the metal composition. Kou [7] describes thatsegregation is important, however, since impurities such as sulfurand phosphorus tend to form low-melting-point films betweensolidifying grains of the metal. These impurities relate to weldjoint detail and the welding process, since they affect the amountof dilution.

Besides, the predominant phase transformation in the mar-tensitic AISI440C stainless steels is austenite-to-martensite thatoccurs in the fusion zone and the as-welded hardness is usuallyin the range of 35–55 HRC with greater cracking risk. Again,the combined influence of grain size and precipitation behaviouron weld metal and HAZ toughness and ductility is analogousto the high-temperature embrittlement phenomenon, which is

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Fig. 1.1. butt welded surface (a) before and (b) after applying the combined preheat-and-postheat welding technique for the line energy, LE¼15 kJ/m.

Fig. 1.2. Schematic diagram to show the variation in incidence angle, A in a plane

tangential to the sample trajectory.

M.M.A. Khan et al. / Optics & Laser Technology 49 (2013) 125–136126

the characteristic in medium chromium alloys like ferriticAISI430F stainless steels. Postweld treatment is, therefore, almostalways required for welding dissimilar martensitic and ferriticstainless steels and used primarily to temper the martensite that,metallurgically, promotes transformation of martensite to ferriteand very fine carbides as stated by Lippold and Kotecki [8]. Thistransformation reduces strength but improves ductility andtoughness. Besides, as described in [9], postweld thermal treat-ments acts to mitigate the high-temperature embrittlement effectby promoting coarsening of the precipitates, thereby reduce theirnegative influence.

The severe heating rate induced by a moving laser source on ametal surface generates a thermal shock having a deep impact onthe change in weld microstructure as well as the formation ofdefects particularly when dissimilar steels are welded. This isbecause the material is brought to the melting and vaporizationstate almost instantaneously along isotherms ruled by the socalled modified Bessel function of the second type and zero orderas reported in [10]. Besides, Huang and Sun [11] state that, ascompared to heating rate, cooling rates can be less dangerous forthe material integrity since temperature decreases with exponen-tial laws offering lower gradients in the tail of the melt pool.Therefore, the idea of the present paper is to generate combinedpre-and-post-heating areas in which the weld material experi-ences less severe heating and cooling rates. This is obtained byusing a non-vertical configuration (A: 151–301–451) of the laserbeam with respect to the material surface as schematically shownin Fig. 1.2.

Laser welding is usually done with the aim of producing anexcellent joint at low cost. However, achieving such a weldjunction without optimization is impossible as stated in [12].Benyounis and Olabi [13] report that various optimization meth-ods are applied to define the desired output variables throughdeveloping mathematical models. In the last two decades, the useof factorial design of experiment (DOE) has grown rapidly and avariety of industries have employed this technique to improveproducts or manufacturing processes as described in [14]. It is apowerful and effective technique to solve challenging qualityproblems. In practice, as stated by Yang et al. [15], the factorialdesign of experiment technique has been used quite successfullyin several industrial applications as in designing electrical/mechanical components or optimizing various manufacturingprocesses.

As seen in the above literatures, weldability is very complexand all the welding factors are interrelated. To better understandthe weldability, it is necessary to study the weld joint andgeometrical features determining the weld shape. The alloycomposition and pre-and-postweld processing must be consid-ered when selecting the welding parameters due to the highcooling and heating rates associated with laser welding. Besides,

no result has been reported yet in the literature on the laser buttwelding of these particular dissimilar ferritic AISI430F and mar-tensitic AISI440C stainless steels or for similar industrial applica-tions. Moreover, in this study, a novel combined laser welding andpre-and-postweld treatment technique is developed for over-coming the problems associated with laser welding of dissimilarstainless steels and applied successfully to the fabrication of fuelinjector. This paper, therefore, examines and optimizes the laserwelding of dissimilar martensitic/ferritic stainless steels in aconstrained butt joint configuration. This study is focused on:

’
Combined welding and pre-and-postweld treatment conceptand its effects,
’
Effects of laser welding parameters such as laser power,welding speed and incidence angle of the beam, impingedtangentially in the direction of rotation as shown in Fig. 1.2, onthe geometrical features of the weld i.e., on weld width,resistance length, and penetration depth,
’
Effects of line energy or energy per unit length on the sameweld geometrical features to understand the energy depen-dent welding phenomena, and

M.M.A. Khan et al. / Optics & Laser Technology 49 (2013) 125–136 127

’
Finally, determination of optimal range of laser weldingparameters, using the developed models with numericaloptimization, to minimize the weld width and maximize theweld penetration depth.
2. Materials and experimental procedures

2.1. Materials

Two cylindrical shells made of ferritic AISI430F (cold drawn,annealed and centerless ground) and martensitic AISI440C (pre-hardened and tempered) stainless steels are welded circularly tomake a butt joint. This dissimilar joint is selected based on bothtechnical and economical aspects, because they can providesatisfactory service performance and considerable savings. More-over, in automotive industries, these materials are frequentlyused in welded form for making different types of fuel injectors.The chemical compositions of base metals available in as-receivedcondition and the weld seam characteristics are shown in Fig. 2.1and Table 2.1, respectively. The inside diameter of the outer shelland the outside diameter of the inner shell are machined toØ7.570.025 mm and Ø7.45870.015 mm, respectively, to have aclearance fit between them when the shells are assembled.

2.2. Experimental procedures

Specimens are welded circularly in a fillet joint configurationusing a 1.1 kW continuous wave Nd:YAG laser (Rofin DY011).The optical system consisted of a 300 mm fiber and two lensesof 200 mm focal and collimate lengths are used to deliver thelaser with a minimum focal spot diameter of 300 mm. A three-stepprocedure is followed to locate the focal point. First, an excep-tionally sharp-nosed tool of 200 mm in height is attached to laser

RmDp

Inner Shell Outer Shell

AISI 440C

AISI 430F

W

RmDp


AISI 440C

AISI 430F

W

RmDp


AISI 440C

AISI 430F

W

Fig. 2.1. Characterization of welding cross-section (W: Weld wid

Table 2.1Chemical compositions of base metals of the weld.

Base metal Composition (in weight percent)

C Cr Ni Mn

AISI430F 0.12 16.0–18.0 0.75 1.0

AISI440C 0.95–1.2 17.2 – 1.00

head mounted on Z motion stage. The laser head is then settangentially in the direction of rotation to an intended beamincident angle. Finally, the positions of the X–Y–Z motion stagesare adjusted in such a way that pointed tool tip touches theplanned point of focus. Laser beam is focused on this located pointthrough the laser head at the specified angle, and the necessaryrotary motion is provided to the specimen through specimenholder mounted on an X–Y motion stage. Computer control panelis interfaced with the linear X–Y–Z as well as rotary motionsystems to regulate the aforesaid movements.

The experiment is initially planned based on statistical factorialexperimental design with full replication. During experimenta-tion, laser power, (P), welding speed (S), and beam incident angle(A) are selected as process input variables for laser welding.Table 2.2 shows the experimental condition, laser welding inputvariables, and design levels used at a glance. Each of the inputvariables and its working range is selected based on industriallyrecommended laser-welding parameters used in automotiveindustries.

General Full Factorial Design is used as a statistical design ofexperiment technique to develop statistical models relating thewelding input parameters to each of the two output responsesof the weld (weld width and penetration depth). The adequaciesof the models developed and their significant linear and interac-tion model terms are measured by analyzing variance and otheradequacy measures. Finally, these mathematical models are usedto determine the optimal settings of welding parameters toensure the desired weld quality. In this study, the quality criteriadefined for the weld to determine the optimal settings of weldingparameters are the minimization of weld width and themaximization of weld penetration depth.

Besides, the energy delivered per unit length of weld line isreferred to as line energy (LE), which is frequently used in variouslaser-processing techniques and termed as a key-parameter whencontinuous-wave laser is used. This term is calculated as the ratio

Outer Shell

Inner Shell

Ø7.458±0.015

Ø9.56±0.03

th, Dp: Weld penetration depth, Rm: Minimum crack-path).

P S Si Mo Se Fe

0.04 0.03 1.00 – – Remainder

– 0.015 1.00 0.75 0.20 Remainder


of laser power over the welding speed as shown in Eq. (1):

LE¼ 0:06�P

SkJ=m� �

ð1Þ

where, LE is line energy; P is laser power in watt (W) describingthe thermal source; and S is welding speed in m/min determiningthe irradiation time. According to the Eq. (1), the combinations oflaser power of 600–1000 W and welding speed of 2.0–4.0 m/minresulted in nominal line energy input in the range of 9.0–30.0 kJ/m.

During experimentation, laser beam is focused on the point ofinterface at a certain angle with the direction of rotation. Argon isused as shielding gas with a constant flow rate of 29 l/min toprotect weld surface from oxidation and suppress the generationof plasma during welding. A standard washing procedure, whichis practised in the automotive industries, is followed to clean, cooland dry the specimens. The experimental set-up for the laser-welding system is illustrated in Fig. 2.2.

SpecimenShielding

Gas Nozzle

Specimen Holder

Laser

Specimen

Head

Fig. 2.2. Photographic view of Nd:YAG laser-welding system.

Table 2.2Experimental conditions and response factors.

Process factors Symbols Actual valuesLaser power (W) P 600

Welding speed (m/min) S 2.0

Angle of incidence (1) A 15

Constant factorsBase material Outer shell

Inner shell

Laser source Continuous wave Nd:YAG laser

Shielding gas Type

Flow rate

Response factorsWeld bead characteristics Weld width (W), weld penetration depth (Dp),

2.3. Weld bead characterization

Welding tests are carried out in a random order to avoid anysystematic error in the experiment. After welding, transversesections are prepared by cutting the samples axially usingSampleMet II (Beuhler, IL) model abrasive cutter. The sectionedsamples are mounted, polished, and etched for mechanicalcharacterization. Software, Leica IM500, incorporated with anoptical microscope (Leica MZ125) is used to measure weld width,penetration depth, and minimum crack path as shown in Fig. 1.Each set of experiments is replicated three times to ensurestatistical accuracy. The mean value of each measured responseparameter is determined and recorded for further analysis.Table 2.3 shows the average measured responses for variouslaser-welding conditions.

The guidance on quality levels for imperfections given in ISO13919-1:1996 is followed to assure the desired weld quality.At this point, each welded specimen is visually inspected before

Table 2.3Design matrix with actual factors and measured mean responses.

Standard order Process factors Response factors

P (W) S (m/min) A (deg) W (mm) Rm (mm) Dp (mm)

1 600 2.0 15 930.7 532.5 599.9

2 800 2-0 15 1050.3 832.8 956.6

3 1000 2.0 15 1021.9 1180.6 1234.5

4 600 3.0 15 639.6 397.9 438.5

5 800 3.0 15 712.5 822.1 833.8

6 1000 3.0 15 851.5 1129.1 1148

7 600 4.0 15 584.6 292.3 301.7

8 800 4.0 15 712.6 563.4 594

9 1000 4.0 15 881.5 618.2 1043.9

10 600 2.0 30 1002.1 375.8 406.2

11 800 2.0 30 1146.8 524.9 593.5

12 1000 2.0 30 1217.7 799.4 959.7

13 600 3.0 30 742.3 195.7 282.8

14 800 3.0 30 897.6 601 603.9

15 1000 3.0 30 899.5 669.2 1074.4

16 600 4.0 30 733.6 249.3 287.4

17 800 4.0 30 820.8 456.1 464.2

18 1000 4.0 30 848 510.7 819.2

19 600 2.0 45 912.8 166.8 256.3

20 800 2.0 45 1024.6 448.1 631.3

21 1000 2.0 45 1073.5 539 801.4

22 600 3.0 45 715.8 174.5 261.9

23 800 3.0 45 843.4 376.6 458

24 1000 3.0 45 933.5 486.1 643.6

25 600 4.0 45 643.3 146.2 205

26 800 4.0 45 707.9 298.2 306.4

27 1000 4.0 45 901.4 413.6 505.3

800 1000

3.0 4.0

30 45

AISI 430F

AISI 440C

Argon

29 l/min

and minimum crack-path (Rm)


and after the cut using the optical microscope. Hermetic weld isensured by performing leak test in vacuum for each of weldedspecimens. During leak test, nitrogen is inflated into theassembled part at a pulsed pressure in the range 10–150 bar forthe expected life cycles. This method also guarantees that theweld will not fail during its service life. In case of failure, theinternal cracks generated during the welding process propagateup to the free surface and N2 leakage is detected by a loss ofvacuum into the chamber.

3. Results and discussion

Various weld profile characteristics are measured with axiallycut specimens using an optical microscope and are recorded forfurther analyses described in the succeeding sections.

Perturbation plots are used to illustrate the effects of indivi-dual process parameter such as laser power (P), welding speed (S),and beam incident angle (A) on geometrical features of the welde.g., weld width, weld penetration depth, and minimum crack-path. Contour plots are used to show the two-factor interactioneffects on the same weld bead geometry.

The line energy is plotted against weld width, weld penetra-tion depth, and minimum crack-path with a view to demonstrate

Tv

Tv

Tm

Ta

Ta

Ta

TaMelt pool

Melt poolSpot

Spot

GF

GL

GL

GB

GF GB

GL

GL

Fig. 3.1. Top views of melt pools and basic draft of thermal gradients (G) in

different directions with: (a) laser beam perpendicular to the surface (b) laser

beam inclined on the tangential plane.

Lase

r inc

iden

t dire

ctio

n

Melt pool

s

Fig. 3.2. (a) Top and cross sectional views of an instant during the welding process, (b)

the effects of energy input on weld profile characteristics, and toexplain different laser welding phenomena as well.

3.1. Combined welding and pre-and-postweld treatment concept

and its effects

The main concept adopted to modify the experimental set uphas been described schematically in Fig. 3.1. When a moving laserbeam hits the surface of a metal in perpendicular direction, anasymmetric melt pool is generated (Fig. 3.1(a)). Heat conductionin the surroundings is not isotropic and generates differentthermal gradients in frontal (GF), lateral (GL) and back (GB)directions. The entity of such difference can be calculated follow-ing [10]. The lengths of the arrows represent a realistic proportionamong the modules of the thermal gradients consideringthe jump between vaporization temperature TV and ambienttemperature TA, and the distance in which it is achieved [3].The frontal gradient is extremely higher with respect to the backdue to the deformation of isotherms caused by the welding speed.The basic idea is to reduce this shock by decreasing GF: this effectis obtained displacing and widening the laser spot (dot line),as reported in Fig. 3.2(b).

From the experimental point of view, this is obtained byinclining the beam on the tangential plane while the specimenmoves against the inclined beam as shown in Fig. 3.2(b). For anygeneric position of the butt specimen welded with an inclinedbeam shown in Fig. 1.2, the cross sectional view of the melt poolis found to be also asymmetric with respect to the beam axis as isreported in Fig. 3.2(a) [16]. This is because the portion of meltpool surface exposed directly to the laser beam is depleted moreas compared to other portion of the keyhole surface.

The top view shows an elliptical print in which the left side iswider than the right one thus generating a thermal gradient oflower entity in the cross sectional plane. This phenomenoncausing the enlargement of the melt pool and the unsymmetricaldistribution of energy inside the print is adopted in the presentresearch to generate a preheating effect before fresh materialreaches the beam axis

During welding, the fresh material is first exposed to the laserradiation in a region with positive defocus which determines thereduction in the thermal gradient. It then reaches the highestintensity zone of the laser beam. At the exit (right) side, the beamacting on welded material results in slight postheat treatmentof the weld. No relevant changes occur in the lateral gradient.The ultimate result is a welded seam due to continuous rotationof the unsymmetrical weld pool over 3601 that reduces the

Focal plane

+ defocus

-defocus

Laser beam

Solidified seamAISI 430FAISI 440C

ample rot. speed

A

draft of the basic concept to obtain a less severe heating rate on the fresh material.

Perturbation

Deviation from Reference Point (Coded Units)-1.00 -0.50 0.00 0.50 1.00

100

375

650

925

1200

P

PS

S

A

A

Min

imum

Cra

ck P

ath,

µm

Perturbation

Deviation from Reference Point (Coded Units)

Wel

d P

enet

ratio

n D

epth

, µm

- 1.00 -0.50 0.00 0.50 1.00

200

475

750

1025

1300

P

P

S

S

A

A

Perturbation

Deviation from Reference Point (Coded Units)

Wel

d W

idth

, µm

-1.00 -0.50 0.00 0.50 1.00

580

740

900

1060

1220

P

P

S

S

AA

Fig. 3.3. Perturbation plot showing effect of all factors on (a) weld width, (b) weld penetration depth, (c) minimum crack path.


occurrence of crack formation as can be seen in Fig. 1.1. Thiscombined process also decreases the penetration depth of theweld since welding with inclined laser beam causes reflectionlosses and a more complex heat transmission through refractiondescribed by Snell’s law.

3.2. Effects of process parameters

The Fig 3.3(a)–(c) show the perturbation plots to compare theeffects of all the process parameters at the center point in thedesign space. The results suggest that laser power has the mostsignificant positive impact on the weld width; weld penetrationdepth; and minimum crack-path. The opposite phenomena areobserved for the welding speed. This is because higher laserpower and slower welding speed result in higher energy deposi-tion on the weld area, and longer irradiation time for thedeposited energy to diffuse into material.

These figures also illustrate that increase in beam incidentangle results in shallower weld penetration, and shorter mini-mum crack-path, whereas the larger beam incident angle causesthe wider weld width. These are due to following consequences:(i) the larger the angle of incidence, the higher the energy loss tothe surrounding through reflection; and (ii) dominance of uni-form heat conduction in all directions over the z-preferential oraxial heat conduction for the lower energy input to the materialsto be welded.

The contour plots shown in Fig. 3.4(a)–(c) demonstrate thefacts that interactions of higher laser power and slower weldingspeed cause wider weld width; deeper weld penetration; andlonger minimum crack-path.

3.3. Effects of line energy

Fig 3.5(a)–(c) show the effects of line energy input on the weldpenetration depth (DP), minimum crack path (Rm), and weldwidth (W), respectively, whereas variation in penetration sizefactor with line energy is illustrated in Fig 3.5(d).

For line energy in the range of 9.0–15 kJ/m, as illustrated inFig 3.5(c)–(d), there is a rapid growth in weld width (W) withenergy input, whereas change in penetration size factor (W/Dp) isinsignificant. Slight negative variations in penetration size factorprove that the laser welding is mainly conduction limited. Sincethe melt pool geometry depends on energy intensity, uniformconduction occurring in all directions usually results in semi-circular weld profile. However, the heat conduction along thebeam axis becomes dominant with the increase in energy inputand weld shape changes from semi-circular to parabolic. Similarparametric effects on the welding pool width at surface [17] andthe upper width [18] are also observed for welding dissimilar lowcarbon and austenitic stainless steels in a butt joint configurationusing the CO2 laser.

Again, a small variation in weld width is observed for the lineenergy input in the range of 15–20 kJ/m. Nonetheless, as shownin Fig 3.5(a) and (d), a sharp decrease (starting from 15 kJ/m) inW/Dp demonstrates the fact that the weld penetration depthincreases at a faster rate than the weld width in this range andestablishes a keyhole formation regime. As a result, the weld beadbecomes almost cylindrical. Penetration size factor increases withfurther increase in line energy. This is due to the creation of upperkeyhole plasma plume that acts as a point heat source above weldplane. This generated plasma plume acts in the keyhole and formsa ‘chalice’ shaped weld bead profile when energy input is more

600 700 800 900 10002.0

2.5

3.0

3.5

4.0Weld Width, µm

Laser Power, W

Wel

ding

Spe

ed, m

/min

791

871

950

1030

1110

737

600 700 800 900 10002.0

2.5

3.0

3.5

4.0Minimum Crack Path, µm

Laser Power, W

Wel

ding

Spe

ed, m

/min

307406 504

603

701

600 700 800 900 10002.0

2.5

3.0

3.5

4.0Weld Penetration Depth, µm

Laser Power, W

Wel

ding

Spe

ed, m

/min

355

487 618 750

882

Fig. 3.4. Contour graphs to show the interaction effects of P and S on (a) weld width, (b) weld penetration depth, and (c) minimum crack path at A¼301.


than 20 kJ/m, which is quite similar to the result obtained from anexperimental study conducted by Khan et al. [5].

Besides, as illustrated in Fig 3.5(b), variation in minimumcrack-path with the line energy input shows the same trend asthe weld penetration depth. This is because of the existing linear,positive relationship between them as can be seen in Fig. 3.6.

3.4. Process parameter optimization

The optimization part in Design-Expert software V7 searchesfor a combination of factor levels that simultaneously satisfy therequirements placed (i.e., optimization criteria) on each of theresponses and process input factors (i.e., multiple-response opti-mization). Numerical and graphical optimization methods areused in this work by selecting the desired goals for each factorand response. As mentioned before, the numerical optimizationprocess involves combining the goals into an overall desirabilityfunction (D). The numerical optimization feature in the design-expert package finds one point or more in the factors domain thatmaximizes this objective function. In this study, the objective is tooptimize the autogenous laser-welded joints subjected to mini-mize the weld width and maximize weld penetration depth acharacterizing factor that determines the minimum crack-path asshown in Fig. 3.6. In order to achieve these objectives, mathema-tical models are developed to relate the aforesaid geometricalfeatures and the selected laser welding input variables.

3.4.1. Development of mathematical models

At this stage, the fit summary in the design-expert software isused to select the models that best describe the response factors.

The fit summary includes sequential model sum squares to selectthe highest order polynomial where additional terms are signifi-cant and the model is not aliased. In addition, model summarystatistics of the fit summary focuses on the model that maximizesadjusted R-squared and predicted R-squared values. The sequen-tial F-test is carried out using the same statistical softwarepackage to check if the regression model is significant and findout the significant model terms of the developed models as well.The step-wise regression method is also applied to eliminate theinsignificant model terms automatically.

Suitable response models for the response factors are selectedbased on the fit summaries. From fit summary output of themeasured responses shown in Tables 3.1–3.4, it is evident thatquadratic model is statistically significant for the weld width,whereas for weld penetration depth; two-factor interaction (2FI)models are statistically recommended for further analyses.

The test for significance of the regression models and the testfor significance on individual model coefficients are performedusing the same statistical package. By selecting the step-wiseregression method that eliminates the insignificant model termsautomatically, the resulting ANOVA Tables 3.5 and 3.6 for theselected models summarize the analysis of variance of eachresponse and illustrate its significant model terms as well. Theaforestated tables demonstrate that calculated Fisher’s ‘Model-F’and ‘Model-P’ values are, respectively, 57.17 and o0.0001 forweld width quadratic model; and 107.74 and o0.0001 for weldpenetration depth 2FI model. These ‘Model-F’ and ‘Model-P’values imply that the selected models are highly significant andthere is only a less than 0.01% chance that these large ‘Model-F’values could occur due to noise. The associated P value is alsoused to estimate whether F is large enough to indicate statistical

5 10 15 20 25 30 35

Wel

d P

enet

ratio

n D

epth

, Dp

(µm

)

0

200

400

600

800

1000

1200

1400

15°30°45°

Line Energy, LE (kJ/m)

15°30°45°

5 10 15 20 25 30 35

Min

imum

Cra

ck P

ath,

Rm

(µm

)

0

200

400

600

800

1000

1200

1400

15°30°45°


15°30°45°

5 10 15 20 25 30 35

Wel

d W

idth

, W (µ

m)

500

600

700

800

900

1000

1100

1200

1300

15°30°45°


15°30°45°

Line Energy, LE (kJ/m)5 10 15 20 25 30 35

Pen

etra

tion

Siz

e Fa

ctor

, W/D

(a.u

.)

0.0

0.5

1.0

1.5

2.0

2.5

3.0

3.5

15°30°

Fig. 3.5. Effect of line energy on (a) weld penetration depth, (b) minimum crack-path, (c) weld width, and (d) penetration size factor for different beam incident angle.

Weld Penetration Depth, Dp (µm)0 200 400 600 800 1000 1200

Min

imum

Cra

ck P

ath,

Rm

(µm

)

100

200

300

400

500

600

700

800

900

30°30°

Fig. 3.6. Relationship between the minimum crack-path and weld penetration

depth.

Table 3.1Sequential model sum of squares for weld width model.

Source Sum ofsquares

df Meansquare

Fvalue

p-valueprob4F

Mean 4.372Eþ007 1 4.372Eþ007

Linear 1.325Eþ006 3 4.417Eþ005 29.28 o0.0001

2FI 3042.62 3 1014.21 0.059 0.9808

Quadratic 2.340Eþ005 3 77994.39 13.15 0.0001 Suggested

Cubic 53388.46 7 7626.92 1.65 0.2363 Aliased

Residual 41493.34 9 4610.37

Total 4.538Eþ007 26 1.745Eþ006

Table 3.2Model summary statistics for weld width model.

Source Std. dev. R2 Adj R2 Pred R2 PRESS

Linear 122.83 0.7997 0.7724 0.7298 4.477Eþ005

2FI 131.56 0.8015 0.7389 0.6687 5.490Eþ005

Quadratic 77.01 0.9427 0.9105 0.8369 2.702Eþ005 Suggested

Cubic 67.90 0.9750 0.9304 0.7707 3.799Eþ005 Aliased


significance. If P value is lower than 0.05, it indicates that themodel is statistically significant as stated by Zulkali et al. [19].

The same tables also show other adequacy measures e.g.,R-squared, adjusted R-squared, and predicted R-squared values. Allthe adequacy measures are in logical agreement and indicatesignificant relationships. Moreover, adequate precision comparesrange of predicted value at the design points to average predictionerror. The adequate precision ratios in all cases are dramaticallygreater than 4 indicating adequate models discrimination.

Again, the ANOVA table for the weld width model shows thatthere is a quadratic relationship between weld width and weldingparameters. The linear terms of laser power (P) and welding speed(S); and the quadratic terms of welding speed and incident angle arethe significant model terms associated with the weld width. How-ever, linear term of beam incident angle is added to support hierarchyof weld width model. For the weld penetration depth model, ANOVAtable demonstrates that all three linear terms i.e., laser power (P),welding speed (S) and beam incident angle; and two-factor


interactions (2FI) of laser power-welding speed (P–S), are the sig-nificant model terms.

From the results shown in Tables 3.1–3.6, it is, therefore,apparent that the developed statistical models for predicting weldwidth and penetration depth are fairly accurate and can be offollowing forms:

(i)

TablSequ

So

Me

Lin

2F

Qu

Cu

Re

To

TablMod

So

Lin

2F

Qu

Cu

Weld width

ðWÞ1:06¼ 2163:33þ0:76P�1154:6Sþ36:97A

þ154:87S2�0:58006A2

(ii)
Weld penetration depth
Dp ¼�446:02þ2:16 P�106:24 Sþ7:674 A�0:024 P x A

e 3.3ential model sum of squares for weld penetration depth model.

urce Sum ofsquares

df Meansquare

Fvalue

p-valueprob4F

an 9.442Eþ006 1 9.442Eþ006

ear 2.046Eþ006 3 6.821Eþ005 95.30 o0.0001

I 61301.75 3 20433.92 4.04 0.0223 Suggested

adratic 8834.15 3 2944.72 0.54 0.6620

bic 50208.79 7 7172.68 1.74 0.2158 Aliased

sidual 37119.61 9 4124.40

tal 1.165Eþ007 26 4.479Eþ005

e 3.4el summary statistics for weld penetration depth model.

urce Std. dev. R2 Adj R2 Pred R2 PRESS

ear 84.60 0.9285 0.9188 0.8994 2.216Eþ005

I 71.14 0.9564 0.9426 0.9240 1.674Eþ005 Suggested

adratic 73.88 0.9604 0.9381 0.9006 2.191Eþ005

bic 64.22 0.9832 0.9532 0.8310 3.725Eþ005 Aliased

Table 3.5ANOVA table for weld width quadratic model.

Source Sum of squares df Mean squar

Model 1.549Eþ006 5 3.097Eþ005

P 3.876Eþ005 1 3.876Eþ005

S 8.518Eþ005 1 8.518Eþ005

A 19093.30 1 19093.30

S2 1.405Eþ005 1 1.405Eþ005

A2 93119.36 1 93119.36

Residual 1.083Eþ005 20 5417.49

Cor total 1.657Eþ006 25

R2¼0.9346 Adj R2

¼0.9183 Pred

Table 3.6ANOVA table for weld penetration depth 2FI model.

Source Sum of squares df Mean squar

Model 2.101Eþ006 4 5.254Eþ005

P 1.388Eþ006 1 1.388Eþ006

S 1.885Eþ005 1 1.885Eþ005

A 4.895Eþ005 1 4.895Eþ005

P–A 55065.56 1 55065.56

Residual 1.024Eþ005 21 4876.13

Cor total 2.204Eþ006 25

R2¼0.9535 Adj R2

¼0.9447 Pred

Normality of residual data and amount of residuals in predic-tion are then checked to ensure statistical validation of thedeveloped models. The normality of data is verified by plottingthe normal probability plot (NPP) of residuals. The residual is thedifference between observed and predicted values (or fittedvalue) obtained from the regression model. The data set isnormally distributed if the points on the plot fall fairly close tothe straight line. The normal probability plots of residual valuesfor weld width, and penetration depth are illustrated in Fig3.7(a)–(b), respectively. The experimental points are reasonablyaligned with predicted or fitted points suggesting the normality ofdata. This is an implication that empirical distribution of residualdata is well-compared with a normal distribution having thesame mean and variance

Fig 3.8(a)–(b) are showing the relationships between theactual and predicted values of weld width and penetration depth.Since the points plotted are close to and around the diagonal line,the difference between the predicted and actual value for eachpoint can be considered to be minimal. It is also an indication thatthe statistical models for prediction are adequate and predictedresults are in good agreement with the measured data.

3.4.2. Numerical optimization

Two criteria are introduced in this numerical optimization.The first set of criteria is to maximize weld penetration depthwith no limitation on either process parameters or weld width. Inthis case, all the process parameters and weld width (firstresponse) are set within a specified range. Furthermore, loweringthe laser power and increasing the welding speed are the mostcommon techniques used in automotive industries to producerelatively low-cost and excellent weld joints. Taking these costand quality aspects into account, second set of criteria for processparameter optimization is fixed to maximize weld penetrationdepth and welding speed, and minimize the laser power and weldwidth. Table 3.7 summarizes these two criteria, lower and upperlimits as well as importance for each input and response factor.

Tables 3.8 and 3.9 show the optimal solution based on the twooptimization criteria as determined by design-expert software.The optimization results clearly demonstrate that, whatever the

e F value p-value prob4F

57.17 o0.0001 Significant

71.54 o0.0001

157.24 o0.0001

3.52 0.0751

25.93 o0.0001

17.19 0.0005

R2¼0.8889 Adeq Precision¼25.958

e F value p-value prob4F

107.74 o0.0001 Significant

284.74 o0.0001

38.65 o0.0001

100.38 o0.0001

11.29 0.0030

R2¼0.9324 Adeq Precision¼36.953

Studentized Residuals

Nor

mal

% P

roba

bilit

y

Normal Plot of Residuals

-1.70 -0.82 0.07 0.95 1.84

1

51020305070809095

99

Studentized Residuals

Nor

mal

% P

roba

bilit

y

Normal Plot of Residuals

-1.98 -0.85 0.27 1.40 2.53

1

5102030

5070809095

99

Fig. 3.7. Normal probability plot for weld (a) width, and (b) penetration depth.

Actual

Pred

icte

d

Predicted vs. Actual

800

1075

1350

1625

1900

856 1108 1360 1613 1865Actual

Pred

icte

dPredicted vs. Actual

100

400

700

1000

1300

124.7 407.6 690.5 973.4 1256.3

Fig. 3.8. Scatter diagrams of weld (a) width, and (b) penetration depth.

Table 3.7Optimization criteria used in this study.

Parameters orresponses

Limits Importance Criterion

Lower Upper First Second

P (W) 600 1000 3 Is in range Minimize

S (m/min) 2.0 4.0 3 Is in range Maximize

A (1) 15 45 3 Is in range Is in range

W (mm) 584 1218 5 Is in range Minimize

DP (mm) 206 1235 5 Maximize Maximize

Table 3.8Optimal solutions as obtained based on first criterion.

Soln no. P (W) S (m/min) F (lm) DP (lm) W (lm)

1 998.6 2.02 15.2 1248.1 1084.6

2 999.3 2.16 15.1 1235.9 1039.7

3 998.9 2.03 15 1250 1078.5

4 997.9 2.00 15 1251.7 1088.6

5 999 2.02 15.5 1243.8 1088

6 999.5 2.04 15.6 1240.4 1082.3

7 998.9 2.03 15.2 1247.4 1081.1

8 998.5 2.14 15.2 1235.3 1048.7

9 999.6 2.18 15 1236 1033.9

10 999.7 2.01 15.8 1240.8 1095


optimization criteria, the beam incident angle has to be itslimiting value of 151 to achieve maximum weld penetrationdepth and minimum weld width. This result also supports thediscussion made earlier on the effect of beam incident angle onthe geometrical features of the weld.

Again, Table 3.8 demonstrating the optimal welding conditionsaccording to the first criterion, it is found that maximum weldpenetration depth of 1250 mm is obtained when laser power andwelding speed are set to their respective highest and lowest limits.However, with acceptable weld penetration depth, the laser powercan be minimized to around 790 W and the highest limit of thewelding speed can be used instead of its lowest limit of 2 m/min.

In this case, as shown in Table 3.9, the weld width and penetrationdepth would be of 665 mm and 706 mm, respectively, which aremuch higher than the prerequisite values for the weld. These resultsalso indicate the fact that laser welding ought to be conduction-limited for this particular joint type and laser-material combinationsin order to obtain the optimal geometrical features of the weld.

Since optimal range of laser power and welding speed selectedbased on second criterion is, respectively, much lower and higherthan that obtained for first set of criterion, any combination of process


parameters of the second optimal set would cause less energy inputto constrained butt joints to be made. This reduced energy input toweld materials would, ultimately, result in less distortion, and

Table 3.9Optimal solutions as obtained based on second criterion.

Soln no. P (W) S (m/min) F (lm) DP (lm) W (lm)

1 792.7 3.95 15.00 676.3 705.4

2 794.9 3.94 15.00 681.3 706.1

3 790.1 3.98 15.00 668.5 705.5

4 795.9 3.99 15.00 678.2 708.7

5 789.2 4.00 15.00 664.6 706.2

6 797.2 3.73 15.00 707.6 702.8

7 826.3 3.99 15.00 731.9 723.9

8 816.9 3.78 15.00 737.6 712.6

9 789.5 3.66 15.00 700.9 699.5

10 810.9 3.63 15.00 742.8 710.3

Table 3.10Visual inspection of weld quality.

Process parameters Visual check

P (W) S (m/min) A (1) Cracks Blow holes Spatter

600 2.0 15 0 0 1

800 2.0 15 1 0 2

1000 2.0 15 0 0 2

600 3.0 15 1 0 1

800 3.0 15 1 0 1

1000 3.0 15 1 0 2

600 4.0 15 0 0 0

800 4.0 15 0 0 0

1000 4.0 15 0 0 1

0¼no defect, 1¼exist but acceptable, 2–3¼not acceptable.

A

Fig. 4.1. (a) typical micrograph of laser welding of AISI440C and AISI430F stainless stee

shell and (c) outer shell.

formation of cracks, blow holes, and spatter with a consequentimprove in the weld quality. The results given in Table 3.10 alsosupport the improvement of weld quality at lower laser power andhigher welding speed.

4. Weld microstructure and microhardness

This study is carried out to investigate the various micro-structures typically formed at the weld zone and various locationsclose to both base metals for a line input of 12.0 kJ/m.The particular sample selected for the metallurgical study is oneof those samples that provide the weld with designed geometry.Fig 4.1(a) shows the micrograph of the weld made of AISI440Cand AISI430F stainless steels.

As shown in Fig 4.1(b) and (c), the microstructures at the innershell are cellular dendritic, whereas columnar dendritic structuresform at the outer shell. These microstructures are a result ofsolidification behavior and subsequent solid-phase transforma-tion controlled by melting ration of two materials to be joinedand weld cooling rate [17]. Since base metals consist of marten-sitic AISI440C and ferritic AISI430F stainless steels, the micro-structures that form in the fusion zone must contain a variety ofcomplex martensite-ferrite structures. Moreover, as shown inFig (c), the intra-granular Cr23C6 carbide formation is evident inthe microstructures forming near the fusion zone boundary. As aresult, at the outer shell, average microhardness of the fusionzone is found to be much higher than that of both the heataffected zone and the base metal as can be seen in Fig. 4.2.

Fig. 4.1(b) shows the microstructures of base metal and HAZ ofthe inner shell, which is pre-hardened and-tempered. Bothmicrostructures contain partly spherodized primary carbide par-ticles in a tempered martensitic matrix. However, microstructures

B

ls, and base metal (BM), heat affected zone (HAZ) and Fusion zone (FZ) of (b) inner

200

300

400

500

600

Local Microhardness

AISI430F

AISI440C

BM BMFZ Boundary

FZBoundary

HAZ HAZ

Welding Pool

Intersection Line

Vick

er's

Har

dnes

s N

umbe

r

Local Microhardness

AISI430F

AISI440C

HAZ HAZ

Intersection Line

Fig. 4.2. Vicker’s local microhardness profile along the line shown in Fig. 4.1(a).


formed in HAZ are finer and contain secondary carbide particlestoo. Formation of these microstructures can be attributed tocombined effects of dissolution of base metal carbide into thesolution in the austenite due to sufficiently high temperature thatprevails in HAZ and repeated hardening due to self-quenching,which occurs inherently during laser welding. As a result, innershell HAZ becomes fully martensitic on cooling, and peak weldhardness occurs in this region as shown in Fig. 4.2. The carbidesformed in this HAZ are normally of M23C6 types, where the ‘M’ ispredominantly of Cr and Fe.

5. Conclusions

Using the laser machine and within the limits of the laserparameters considered in this study the following points can beconcluded:

’
The developed combined welding and pre-and-postweldtreatment technique is able to overcome the crack formationproblem associated with laser welding of dissimilar stainlesssteels.
’
Laser power and welding speed are the most significant laserwelding input factors and have opposite effects.
’
For welding dissimilar stainless steels in a butt joint config-uration, weld penetration depth determines the minimumcrack-path of the weld.
’
Formation of keyhole results in rapid change in the weldgeometrical features within a certain range of energy input. Afterthe upper limiting value, formation of upper keyhole plasmaplume only contributes to the change in shape of the weld bead.
’
Various, complex martensite–ferrite microstructures developin the fusion zone. As a consequence, average microhardnessof the fusion zone becomes much higher than that of basemetal of the outer shell and lower than that of base metal ofthe inner shell.
’
By means of design of experiment inspired by full factorialdesign, it is possible to achieve the best operating parameterwindow.
’
For these particular joint type and laser-material combinations: | A laser power of 790–810 W and a welding speed of 3.6–4.0
m/min are the optimal input parameters to obtain an excellentweld.

|
Whatever the optimization criteria, the beam incident anglehas to be around its limiting value of 151.
’
Efficient and low-cost weld joints could be achieved using thewelding conditions drawn from the numerical optimization
Acknowledgements

Authors would like to acknowledge Continental ‘‘Cleanlinessand Material Lab’’ for its support while cutting and preparing thesamples for analysis. Thanks are due to M. Fiaschi and F. Sarri,Industrial Engineering Department for their kind assistance atvarious stages of this research. The help extended by V. Colombiniand C. Fierro during experimental investigation is also sincerelyacknowledged.

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