50363_02a

2Chemical Reactions in Arc Welding

2.1 Introduction

The weld metal composition is controlled by chemical reactions occurring in the weld pool atelevated temperatures, and is therefore influenced by the choice of welding consumables (i.e.combination of filler metal, flux, and/or shielding gas), the base metal chemistry, as well as theoperational conditions applied. In contrast to ladle refining of metals and alloys where thereactions occur under approximately isothermal conditions, a characteristic feature of the arcwelding process is that the chemical interactions between the liquid metal and its surroundings(arc atmosphere, slag) take place within seconds in a small volume where the metal tempera-ture gradients are of the order of 1000°C mm"1 with corresponding cooling rates up to1000°C s"1. The complex thermal cycle experienced by the liquid metal during transfer fromthe electrode tip to the weld pool in GMA welding of steel is shown schematically in Fig. 2.1.

As a result of this strong non-isothermal behaviour, it is very difficult to elucidate thereaction sequences during all stages of the process. Consequently, a complete understandingof the major controlling factors is still missing, which implies that fundamentally based pre-dictions of the final weld metal chemical composition are limited. Additional problems resultfrom the lack of adequate thermodynamic data for the complex slag-metal reaction systemsinvolved. However, within these restrictions, the development of weld metal compositionscan be treated with the basic principles of thermodynamics and kinetic theory considered inthe following sections.

2.2 Overall Reaction Model

The symbols and units used throughout this chapter are defined in Appendix 2.1.In ladle refining of metals and alloys, the reaction kinetics are usually controlled by mass

transfer between the liquid metal and its surroundings (slag or ambient atmosphere). Exam-ples of such kinetically controlled processes are separation of non-metallic inclusions from adeoxidised steel melt or removal of hydrogen from liquid aluminium. In welding, the reactionpattern is more difficult to assess because of the characteristic non-isothermal behaviour of theprocess (see Fig. 2.1). Nevertheless, experience shows that it is possible to analyse masstransfer in welding analogous to that in ladle refining by considering a simple two-stage reac-tion model, which assumes:1

(i) A high temperature stage, where the reactions approach a state of localpseudo-equilibrium.

(ii) A cooling stage, where the concentrations established during the initial stagetend to readjust by rejection of dissolved elements from the liquid.

Fig. 2.1. Schematic diagram showing the main process stages in GMA welding. Characteristic averagetemperature ranges at each stage are indicated by values in parenthesis.

As indicated in Fig. 2.2 the high temperature stage comprises both gas/metal and slag/metalinteractions occurring at the electrode tip, in the arc plasma, or in the hot part of the weld pool,and is characterised by extensive absorption of elements into the liquid metal. During thesubsequent stage of cooling following the passage of the arc, a supersaturation rapidly in-creases because of the decrease in the element solubility with decreasing temperatures. Thesystem will respond to this supersaturation by rejection of dissolved elements from the liquid,either through a gas/metal reaction (desorption) or by precipitation of new phases. In the lattercase the extent of mass transfer is determined by the separation rate of the reaction products inthe weld pool. It should be noted that the boundary between the two stages is not sharp, whichmeans that phase separation may proceed simultaneously with absorption in the hot part of theweld pool.

In the following sections, the chemistry of arc welding will be discussed in the light of thistwo-stage reaction model.

2.3 Dissociation of Gases in the Arc Column

As shown in Table 2.1, gases such as hydrogen, nitrogen, oxygen, and carbon dioxide will bewidely dissociated in the arc column because of the high temperatures involved (the arc plasmatemperature is typically of the order of 10 0000C or higher). From a thermodynamic stand-point, dissociation can be treated as gaseous chemical reactions, where the concentrations ofthe reactants are equal to their respective partial pressures. Hence, for dissociation of diatomicgases, we may write:

where X denotes any gaseous species.

(2-1)

Gas nozzleShielding gasFiller wire

Electrode tip droplet(1600-20000C)

Falling droplet (24000C)

Hot part of weld pool(1900-22000C)

Contact tube

Arc plasmatemperature~10000°C

Cold part ofweld pool (< 19000C)

Base plateWeld poolretention

time 2-1Os

Tem

pera

ture

Solid

wel

d m

etal

Solid

wel

d m

etal

Solid

wel

d m

etal

Conc

entra

tion

Solid

wel

d m

etal

'Hot1 part ofweld pool 'Cold' part of weld pool

Peak temperature

Grey jzonei

Rejection ofdissolved elements

Absorptionof elements

Peak concentration

Equilibrium concentration at melting point

Time

Fig. 2.2. Idealised two-stage reaction model for arc welding (schematic).

Table 2.1 Temperature for 90% dissociation of some gases in the arc column. Data from Lancaster.2

Next, consider a shielding gas which consists of two components, i.e. one inert component(argon or helium) and one active component X2. When the fraction dissociated is close tounity, the partial pressure of species X in the gas phase px is equal to:

Gas Dissociation Temperature (K)

CO2 3800

H2 4575

O2 5100

N2 8300

where H1 and nx are the total number of moles of components / (inert gas) and X, respectivelyin the shielding gas, andptot is the total pressure (in atm).

It follows from equation (2-1) that two moles of X form from each mole of X2 that dissoci-ates. Hence, equation (2-2) can be rewritten as:

(2-2)

(2-3)

where nXl is the total number of moles of component X2 which originally was present in theshielding gas.

If nXl and H1 are proportional to the volume concentrations of the respective gas compo-nents in the shielding gas, equation (2-3) becomes:

(2-4)

Similarly, if X2 is replaced by another gas component of the type YX2, we get:

(2-5)

Taking vol% / = (100 - vol% X2) andp,ot = 1 atm, we obtain the following expression for

Px-

(2-6)

(2-7)

and

It is evident from the graphical representations of equations (2-5) and (2-7) in Fig. 2.3 thatthe partial pressure of the dissociated component X increases monotonically with increasingconcentrations of X2 and YX2 in the shielding gas. The observed non-linear variation of px

arises from the associated change in the total number of moles of constituent species in the gasphase due to the dissociation reaction. Moreover, it is interesting to note that the partial pres-sure px is also dependent on the nature of the active gas component in the arc column (i.e. thestoichiometry of the reaction). This means that the oxidation capacity of for instance CO2 isonly half that of O2 when comparison is made on the basis of equal concentrations in theshielding gas (to be discussed later).

Px

Vol%)^f VoRGYX2

Fig. 2.3. Graphical representation of equations (2-5) and (2-7).

2.4 Kinetics of Gas Absorption

In general, mass transfer between a gas phase and a melt involves:3

(i) Transport of reactants from the bulk phase to the gas/metal interface.

(ii) Chemical reaction at the interface.

(iii) Transport of dissolved elements from the interface to the bulk of the metal.

2.4.1 Thin film model

In cases where the rate of element absorption is controlled by a transport mechanism in the gasphase (step one), it is a reasonable approximation to assume that all resistance to mass transferis confined to a stagnant layer of thickness 8 (in mm) adjacent to the metal surface, as shown inFig. 2.4. Under such conditions, the overall mass transfer coefficient is given by:2

(2-8)

where Dx is the diffusion coefficient of the transferring species X (in mm2 s~*).Although the validity of equation (2-8) may be questioned, the thin film model provides a

simple physical picture of the resistance to mass transfer during gas absorption.

Par

tial

pres

sure

Distance

Fig. 2.4. Film model for mass transfer (schematic).

2.4.2 Rate of element absorption

Referring to Fig. 2.5, the rate of mass transfer between the two phases (in mol s"1) can bewritten as:

(2-9)

where A is the contact area (in mm2), R is the universal gas constant (in mm3 atm K"1 mol"1),T is the absolute temperature (in K), px is the partial pressure of the dissociated species X in thebulk phase (in atm), and px is the equilibrium partial pressure of the same species at the gas/metal interface (in atm).

Based on equation (2-9) it is possible to calculate the transient concentration of element Xin the hot part of the weld pool. Let m denote the total mass of liquid weld metal entering/leaving the reaction zone per unit time (in g s"1). If Mx represents the atomic weight of theelement (in g mol"1), we obtain the following relation when /? x » p°x:

(2-10)

It follows from equation (2-10) that the transient concentration of element X in the hot partof the weld pool is proportional to the partial pressure of the dissociated component X in theplasma gas. Since this partial pressure is related to the initial content of the molecular speciesX2 or YX2 in the shielding gas through equations (2-5) and (2-7), we may write:

Arc columnBulk gasphase

Stagnant gaseousboundary layer

Gas/metal interface

Metal phase

Hot part ofweld pool

Fig. 2.5. Idealised kinetic model for gas absorption in arc welding (schematic).

(2-11)

(2-12)

and

where C1 and C2 are kinetic constants which are characteristic of the reaction systems underconsideration.

2.5 The Concept of Pseudo-Equilibrium

Although the above analysis presupposes that the element absorption is controlled by atransport mechanism in the gas phase, the transient concentration of the active component X inthe hot part of the weld pool can alternatively be calculated from chemical thermodynamics byconsidering the following reaction:

(2-13)X(gas) X (dissolved)

By introducing the equilibrium constant K{ for the reaction and setting the activity coeffi-cient to unity, we get:

(2-14)

This equation should be compared with equation (2-10) which predicts a linear relationship

between wt% X and px. If the above analysis is correct, one would expect that the partialpressure px at the gas/metal interface is directly proportional to the partial pressure of thedissociated component in the bulk phase. Unfortunately, the proportionality constant is diffi-cult to establish in practice.

2.6 Kinetics of Gas Desorption

During the subsequent stage of cooling following the passage of the arc, the concentrationsestablished at elevated temperatures will tend to readjust by rejection of dissolved elementsfrom the liquid. When it comes to gases such as hydrogen and nitrogen, this occurs through adesorption mechanism, where the driving force for the reaction is provided by the decrease inthe element solubility with decreasing metal temperatures.

2.6.1 Rate of element desorption

Consider a melt which first is brought in equilibrium with a monoatomic gas of partial pres-sure px at a high temperature T1, and then is rapidly cooled to a lower temperature T2 andimmediately brought in contact with diatomic X2 of partial pressure pXl (see Fig. 2.6). Undersuch conditions, the rate of element desorption (in mol s"1) is given by:

(2-15)

where k'd is the mass transfer coefficient (in mm s 1X and p°x is the equilibrium partial pres-

sure of component X2 at the gas/metal interface (in atm).

Bulk gasphase

Stagnant gaseousboundary layer

Gas/metal interface

Metal phase

Cold part ofweld pool

Fig. 2.6. Idealised kinetic model for gas desorption in arc welding (schematic).

The partial pressure pX2 can be calculated from chemical thermodynamics by considering

the following reaction:2X(dissolved) = X2 (gas) (2-16)

from which(2-17)

where K2 is the equilibrium constant, and [wt% X] is the concentration of element X in theliquid metal (in weight percent). Note that the activity coefficient has been set to unity in thederivation of equation (2-17).

The equilibrium constant K2 may be expressed in terms of the solubility of element X in theliquid metal at 1 atm total pressure Sx. Hence, equation (2-17) transforms to:

(2-18)

(2-19)

By combining equations (2-15) and (2-18), we get:

Data for the solubility of hydrogen and nitrogen in some metals up to about 22000C aregiven in Figs. 2.7 and 2.8, respectively. It is evident that the element solubility decreasessteadily with decreasing metal temperatures down to the melting point. This implies that thedesorption reaction is thermodynamically favoured by the thermal conditions existing in thecold part of the weld pool.

2.6.2. Sievert's law

It follows from equation (2-19) that desorption becomes kinetically unfeasible whenPx2 ~ Px2' corresponding to:

(2-20)

Equation (2-20) is known as the Sievert's law. This relation provides a basis for calculatingthe final weld metal composition in cases where the resistance to mass transfer is sufficientlysmall to maintain full chemical equilibrium between the liquid metal and the ambient (bulk)gas phase.

2.7 Overall Kinetic Model for Mass Transfer during Cooling in the Weld Pool

Because of the complexity of the rate phenomena involved, it would be a formidable task toderive a complete kinetic model for mass transfer in arc welding from first principles. How-

ml H

2/100

g fu

sed

met

al

ml H

2/100

g fu

sed

met

al

ml H

2/100

g fu

sed

met

al

ml H

2/100

g fu

sed

met

al

(a) (b)

Aluminium

Temperature, 0C

Copper

Solid Cu

Temperature, 0C

(C) (d)

Iron Nickel

Temperature, 0CTemperature, 0C

Fig. 2.7. Solubility of hydrogen in some metals; (a) Aluminium, (b) Copper, (c) Iron, (d) Nickel. Datacompiled by Christensen.4

ever, for the idealised system considered in Fig. 2.9, it is possible to develop a simple math-ematical relation which provides quantitative information about the extent of element transferoccurring during cooling in the weld pool. Let [%X]eq denote the equilibrium concentration ofelement X in the melt. If we assume that the net flux of element X passing through the phaseboundary A per unit time is proportional to the difference ([%X] - [%X]eqX the followingbalance is obtained:3

where V is the volume of the melt (in mm3), kd is the overall mass transfer coefficient (inmm s"1), and A is the contact area between the two phases (in mm2).

(2-21)

log (w

t% N

)

Net f

lux

of X

Dist

ance

Temperature, 0C

Iron

104AT1 K

Fig. 2.8. Solubility of nitrogen in iron. Data from Turkdogan.5

Phase I i

Contact area (A)

Phase i Volume (V)

Concentration

Fig. 2.9. Idealised kinetic model for mass transfer in arc welding (schematic).

By rearranging equation (2-21) and integrating between the limife [%X]( (att = O) and [%X](at an arbitrary time t\ we get:

(2-22)

where to is a time constant (equal to VI kjA).

(X-X

^)Z(

X1-X

eq)

Fig. 2.10. Graphical representation of equation (2-22).

t , s

Under such conditions the final weld metal composition can be calculated from simplechemical thermodynamics.

Because of this flexibility, equation (2-23) is applicable to a wide range of metallurgicalproblems at the same time as it provides a simple physical picture of the resistance to masstransfer during cooling in the weld pool.

(2-24)

It follows that the final concentration of element X in the weld metal depends both on thecooling conditions and on the intrinsic resistance to mass transfer, combined in the ratio t/to.When [%X]eq is sufficiently small, equation (2-23) predicts a direct proportionality between[%X] and [%X\t (i.e. the initial concentration of element X in the weld pool). This will be thecase during deoxidation of steel weld metals where separation of oxide inclusions from theweld pool is the rate controlling step. Moreover, when t/t0 » 1 (small resistance to masstransfer), equation (2-23) reduces to:

(2-23)

It is evident from the graphical representation of equation (2-22) in Fig. 2.10 that the rate ofmass transfer depends on the ratio Vl kji, i.e. the time required to reduce the concentration ofelement X to a certain level is inversely proportional to the mass transfer coefficient kd. Thistype of response is typical of a first order kinetic reaction.

Although the above model refers to mass transfer under isothermal conditions, it is alsoapplicable to welding if we assume that the weld cooling cycle can be replaced by an equiva-lent isothermal hold-up at a chosen reference temperature. Thus, by rearranging equation (2-22), we get:

2.8 Absorption of Hydrogen

Some of the well-known harmful effects of hydrogen discussed in Chapters 3 and 7 (i.e. weldporosity and HAZ cold cracking) are closely related to the local concentration of hydrogenestablished in the weld pool at elevated temperatures due to chemical interactions betweenthe liquid metal and its surroundings.

2.8.1 Sources of hydrogen

Broadly speaking, the principal sources of hydrogen in welding consumables are:6

(i) Loosely bound moisture in the coating of shielded metal arc (SMA) electrodes and inthe flux used in submerged arc (SA) or flux-cored arc (FCA) welding. Occasionally, moisturemay also be introduced through the shielding gas in gas metal arc (GMA) and gas tungsten arc(GTA) welding.

(ii) Firmly bound water in the electrode coating or the welding flux. This can be in theform of hydrated oxides (e.g. rust on the surface of electrode wires and iron powder), hydro-carbons (in cellulose), or crystal water (bound in clay, astbestos, binder etc.).

(iii) Oil, dirt and grease, either on the surface of the work piece itself, or trapped in thesurface layers of welding wires and electrode cored wires.

It is evident from Fig. 2.11 that the weld metal hydrogen content may vary strongly fromone process to another. The lowest hydrogen levels are usually obtained with the use of low-moisture basic electrodes or GMA welding with solid wires. Submerged arc welding and flux-cored arc welding, on the other hand, may give high or low concentrations of hydrogen in theweld metal, depending on the flux quality and the operational conditions applied (note that theformer process is not included in Fig. 2.11). The highest hydrogen levels are normally associ-ated with cellulosic, acid, and rutile type electrodes. This is due to the presence of largeamounts of asbestos, clay and other hydrogen-containing compounds in the electrode coating.

Table 2.2 (shown on page 132) gives a summary of measured arc atmosphere compositionsin GMA and SMA welding. Included are also typical ranges for the weld metal hydrogencontent.

2.8.2 Methods of hydrogen determination in steel welds

Hydrogen is unlike other elements in weld metal in that it diffuses rapidly at normal roomtemperatures, and hence, some of it may be lost before an analysis can be made. This, coupledwith the fact that the concentrations to be measured are usually at the parts per million level,means that special sampling and analysis procedures are needed. In order that research resultsmay be compared between different laboratories and can be used to develop hydrogen controlprocedures, some international standardisation of these sampling and analysis methods is nec-essary.

Three methods are currently being used, as defined in the following standards:

Pote

ntia

l hyd

roge

n le

vel

FCAW

Verylow Low Medium High

Weld hydrogen level

Fig. 2.11. Ranking of different welding processes in terms of hydrogen level (schematic). The diagram isbased on the ideas of Coe.6

(i) The Japanese method (JIS Z 313-1975), which has been adopted with important ad-justments from the former ASTM designation A316-48T. This method involves collection ofreleased hydrogen from a single pass weld above glycerine for 48h at 45 0C. The total volumeof hydrogen is reported in ml per 10Og deposit. Only 5 s of delay are allowed from extinctionof the arc to quenching.

(ii) The French method (N.F.A. 81-305-1975) where two beads are deposited onto corewires placed in a copper mould. Hydrogen released from this bead is collected above mercury,and the volume is reported in ml per 10Og fused metal (including the fused core wire metal).

(iii) The International Institute of Welding (HW) method (ISO 3690-1977), where a singlebead is deposited on previously degassed and weighed mild steel blocks clamped in a quick-release copper fixture. The weldment is quenched and refrigerated according to a rigorouslyspecified time schedule. Hydrogen released from the specimens is collected above mercuryfor 72 h at 25°C, and the results are reported in ml per 10Og deposit, or in g per ton fused metal.To avoid confusion, it is recommended to use the symbol HDM for the content reported in termsof deposited metal (ml per 10Og deposit), and HFM for the content referred to fused metal (mlper 100 g or g per ton fused metal). The relationship between HDM and HFM is shown in Fig.2.12.

As would be expected, these three methods do not give identical results when applied to agiven electrode. Approximate correlations have been established between the HW criteriaHDM and HFM and the numbers obtained by the Japanese and the French methods (designatedHJIS and HFR, respectively). For covered electrodes tested at various hydrogen levels, wehave:7

The conversion factor from HFR to HFM applies to a ratio of deposited to fused metal,DI(B + D), equal to 0.6, which is a reasonable average for basic electrodes.

The use of HFM in preference of HDM is normally recommended, because it is a more ra-tional criterion of concentration. Moreover, HDM values would be grossly unfair, if applied tohigh penetration processes like submerged arc welding. In GTA welds made without fillerwire HDM cannot be used at all, since there is no deposit.

It should be noted that the present HW procedure gives the amount of 'diffusible hydrogen'.For certain purposes the total hydrogen content may be wanted. It is obtained by adding thecontent of 'residual hydrogen' determined on the same samples by vacuum or carrier gasextraction at 6500C. A very small additional amount may be observed on vacuum fusion of thesample, tentatively labelled 'fixed hydrogen'. There is no clear line of demarcation betweenthese categories of hydrogen. As will be discussed later, the extent of hydrogen trappingdepends both on the weld metal constitution and the thermal history of the metal. In single-bead basic electrode deposits the diffusible fraction is usually well above 90%.

2.8.3 Reaction model

Normally, measurements of hydrogen in weld metals are carried out on samples from solidi-fied beads. Due to the rapid migration of hydrogen at elevated temperatures, such data do notrepresent the conditions in the hot part of the weld pool. Quenched end crater samples wouldbe better in this respect, but they are not representative of normal welding. Further complica-tions arise from the presence of hydrogen in different states (e.g. diffusible or residual hydro-gen) and the lack of consistent sampling methods.

Nevertheless, experience has shown that pick-up of hydrogen in arc welding can be inter-preted on the basis of the simple model outlined in Fig. 2.13. According to this model, twozones are considered:

(i) An inner zone of very high temperatures which is characterised by absorption of atomichydrogen from the surrounding arc atmosphere.

(2-25)

(2-26)

Fig. 2.12. The relation between HDM and HFM (0.9 is the conversion factor from ml per 10Og to g per ton).

Fig. 2.13. Idealised reaction model for hydrogen pick-up in arc welding.

(ii) An outer zone of lower temperatures where the resistance to hydrogen desorption issufficiently small to maintain full chemical equilibrium between the liquid weld metal and theambient (bulk) gas phase.

Under such conditions, the final weld metal hydrogen content should be proportional to thesquare root of the initial partial pressure of diatomic hydrogen in the shielding gas, in agree-ment with Sievert's law (equation (2-20)).

2.8.4 Comparison between measured and predicted hydrogen contents

It is evident from the data in Table 2.2 that the reported ranges for hydrogen contents in steelweld metals are quite wide, and therefore not suitable for a direct comparison of predictionwith measurement. For such purposes, the welding conditions and consumables must be moreprecisely defined.

2.8.4.1 Gas-shielded weldingIn GTA and GMA welds the hydrogen content is usually too low to make a direct comparisonbetween theory and experiments. An exception is welding under controlled laboratory condi-tions where the hydrogen content in the shielding gas can be varied within relatively widelimits. The results from such experiments are summarised in Fig. 2.14, from which it is seenthat Sievert's law indeed is valid. A closer inspection of the data reveals that the weld metalhydrogen content falls within the range calculated for chemical equilibrium at 1550 and 20000C,depending on the applied welding current. This shows that the effective reaction temperatureis sensitive to variations in the operational conditions.

An interesting effect of oxygen on the weld metal hydrogen content has been reported byMatsuda et al.9 Their data are reproduced in Fig. 2.15. It is evident that the hydrogen level issignificantly higher in the presence of oxygen. This is probably due to the formation of a thin(protective) layer of slag on the top of the bead, which kinetically suppresses the desorption ofhydrogen during cooling.

Hot part of weld poolAbsorption of atomic hydrogen

(controlled by pH in the arc column)

Electrode

Cold part of weld poolDesorption of hydrogen

(controlled by pH2 in ambient gas phase)

Hydrogentrapped inweld metal

Weld pool

ml H

2/10

0 g

fuse

d m

etal

, H

pM

Table 2.2 Measured arc atmosphere compositions in steel welding. Also included are typical ranges forthe weld metal hydrogen content. Data compiled by Christensen.4

Method

GMAW*(CO2)

SMAW(acid)

SMAW(rutile)

SMAW(basic)

FCAW(rutile)

FCAW(basic)

SAW(basic)

Primary Sourceof Hydrogen

Moisture introducedthrough the shielding gas

Firmly bound water inthe electrode coating

Firmly bound water inthe electrode coating

Loosely bound water inthe electrode coating

Firmly bound waterinflux

Loosely bound waterinflux

Loosely bound waterinflux

Arc Atmosphere Composition(vol%)

CO2

98-80

-4

~4

-19

CO

2-20

-34

-42

-77

H2+H2O

<0.02

-62

-54

-4

Weld MetalHydrogen

Content (ppm)

Range

1-5

10-30

10-30

2-10

10-20

2-5

2-10

Average

3

25

25

3-5

GTAW (low-alloy steel)

*The arc atmosphere composition can vary within wide limits, depending on the operational conditions applied.

ml H

2 /1

00 g

fuse

d m

etal

, HJ|

SGTAW (low-alloy steel)Welding conditions: 300A-18V-2.5 mm/s

Weld metal oxygen content, wt%

Fig. 2.15. Hydrogen pick-up in GTA welding at different levels of oxygen in the weld metal. Data fromMatsuda et al.9

Example (2.1)

Consider GTA welding (Ar-shielding) on a thick plate of low-alloy steel under the followingconditions:

/ = 200A, U = 15V, v = 3 mm s"1, TI = 0.5, T0 = 20°C

The shielding gas contains 0.1 vol% moisture (H2O) and is supplied at a rate of 15Nl mhr1.Calculate the 'potential' hydrogen level, assuming that all hydrogen introduced through theshielding gas is absorbed in the weld metal.

SolutionFirst we calculate the total mass of hydrogen per mm:

The resulting bead cross section and total mass of weld metal per mm can be estimatedfrom the Rosenthal equation by considering the dimensionless operating parameter at the meltingpoint (equation (1-50)):

Reading from Fig. 1.21 gives:

Taking the density of the steel equal to 7.85 X 10 3 g mm 3, we obtain:

The 'potential' hydrogen level is thus:

It is evident from the above calculations that the 'potential' hydrogen level is at least oneorder of magnitude higher than the expected weld metal hydrogen content (1 to 3 ppm). Thisshows that the hydrogen pick-up in GTA welding is not determined by the total amount ofhydrogen which is introduced through the shielding gas, but is mainly controlled by the result-ing partial pressure of hydrogen in the ambient (bulk) gas phase.

2.8.4.2 Covered electrodesIn SMA welding the partial pressure of hydrogen is more difficult to assess due to the presenceof trapped moisture and hydrogen-containing compounds in the electrode coating. Such com-pounds will loose their identity at the stage of introduction into the arc atmosphere. Since verylittle information is available on the species present in the arc column, we shall base our esti-mate on a simple thermodynamic approach, including only the molecular species H2 and H2Owhich can be determined by analysis (see data in Table 2.2). It follows that the combinedpartial pressure of H2 and H2O in the gas phase is given by:

The parameter pw can be estimated on the basis of combustion measurements of the elec-trode coating, assuming that no carbon is picked up or lost from the system in excess of theamount calculated from an analysis of the base plate and the electrode wire. For a recordedcontent of mw g H2O and mc g CO2 per 100 g of electrode coating, we obtain:

From a thermodynamic standpoint, replacement of pHl by /^ in the expression for Sievert'slaw requires the use of a modified solubility of hydrogen, defined as:

(2-27)

(2-28)

(2-29)

where K3 is the equilibrium constant for the H2O-H reaction, and [%O] is the weld metaloxygen content. In practice, the correction term ^ K3/(K3+[%0]) does not depart signifi-cantly from unity, which means that Sw ~ SH.

ml H

2/100

g fu

sed

met

al, H

pM •

During welding with basic covered electrodes considerable amounts of CO2 may form as aresult of decomposition of calcium carbonate, according to the reaction:

(2-30)

Modern basic electrodes contain between 20 to 40 weight percent CaCO3, which is equiva-lent with a CO2 content of 9 to 18 percent. Taking as an average mc equal to 15 g CO2 per 100 gelectrode coating, we obtain:

(2-31)

In Fig. 2.16 the validity of equation (2-31) has been checked against relevant literature data(compiled by Chew10). A closer inspection of the data reveals that the weld metal hydrogencontent falls within the range calculated for chemical equilibrium at 1520 to 2000°C, taking Sw

equal to the solubility of hydrogen in pure iron at the indicated temperatures (i.e. 27 and 40 mlH2 per 100 g fused metal, respectively). Although the observed scatter in the effective reactiontemperature is admittedly large, equation (2-31) points out a very interesting effect, namelythat the hydrogen content of SMA steel weld metals is controlled by the combined partialpressure of H2 and H2O in the ambient gas phase. For this reason it is frequently recom-mended that calcium carbonate is added to the electrode coating, which on decompositionproduces considerable amounts of shielding gas in the form of CO2. Hydrogen shielding canalso be achieved by additions of volatile alkali-fluorides, which on heating will evaporate anddilute the atmosphere with respect to hydrogen.

SMAW (low-alloy steel)

Fig. 2.16. Hydrogen pick-up in SMA welding at different water contents in the electrode coating. Datacompiled by Chew.10

Water content in electrode coating, wt%

Example (2.2)

Consider SMA welding on mild steel with basic covered electrodes. The electrode coatingcontains 35 wt% CaCO3 and 0.5 wt% H2O in the as-received condition. After drying at 3500Cfor 1 h the water content is reduced to 0.2 wt% H2O. Estimate the weld metal hydrogencontent (in ppm) both before and after drying of the electrode. Assume in these calculations aneffective reaction temperature of 18000C.

SolutionFirst we calculate the CO2 content per 100 g of electrode coating. Taking the atomic weight ofCaCO3 and CO2 equal to 100.1 and 40.0, respectively, we obtain:

The combined partial pressure pw can now be estimated from equation (2-28). Beforedrying we have:

After drying of the electrode, the partial pressure pw becomes:

From Fig. 2.7(c) it is evident that the solubility of hydrogen in liquid iron at 18000C isabout 37 ml H2 per 100 g fused metal. This corresponds to a modified solubility Sw (in ppm)of:

Substituting this value into the expression for Sievert's law gives:

(before drying)(after drying)

It follows from the above calculations that a low weld metal hydrogen level requires the useof 'dry' basic electrodes. In practice, this can be achieved by protecting the electrodes againstmoisture pick-up during storage (see Fig. 2.17). However, in certain cases it is necessary todifferentiate between strongly bound and loosely adsorbed moisture in the coating of basicelectrodes. This point is more clearly illustrated in Fig. 2.18, which shows the HDM content ofhydrogen in basic electrode deposits at various levels of coating moisture. It is seen that waterremaining from an insufficient baking treatment is more dangerous than moisture picked up byexposure of a properly dried coating. This has to do with the fact that loosely adsorbed mois-

ml

H2 /

100 d

eposi

t, H

DM

Wat

er c

onte

nt in

ele

ctro

de c

oatin

g, w

t%

very

low

low

med

ium

high

Fig. 2.18. Hydrogen pick-up in SMA welding at different levels (states) of adsorbed water in the elec-trode coating. Data from Evans and Bach.12

Water content, wt%

SMAW (low-alloy steel)

Fig. 2.17. Moisture content in basic electrode coating as a function of exposure time and relative humid-ity (R.H.) in ambient gas phase. Data from Evans.11

Exposure time, days

ml H

2/10

0 de

posi

t, H

DM

Hyd

roge

n co

nten

t, H

FM (p

pm)

ture will tend to evaporate during the welding operation (before it enters the arc column)because of resistance heating of the electrode, a process which is not feasible when the water isbound in rust on the surface of the electrode wire or the iron powder.

2.8.4.3 Submerged arc weldingThis method is usually classified as a pure slag-shielded process, because carbonates or othergas-producing compounds are not present in large quantities. A closed arc cavity does exist,however, as indicated by the falling volt-ampere curve characteristic of open arcs, and byobservations made by probes inserted through the flux cover.

It is reasonable to assume that the gas contained within this enclosure consists of metalvapour, volatile constituents originating from the flux, and relatively small fractions of carbonmonoxide and water vapour. Acid fluxes of the calcium silicate type will probably generatesilicon monoxide, while agglomerated fluxes bonded with alkali silicate will produce volatilealkali fluorides. In addition, carbon monoxide may be present as a result of oxidation of car-bon, or decomposition of carbonates.

A small but important contribution to the cavity atmosphere is the trace of moisture remain-ing in the flux even after careful drying. No direct measurements of partial pressures are avail-able, and the gas composition must therefore be inferred from observations of hydrogen ab-sorption in the weld metal. Hydrogen pick-up during SA welding has been examined by Evansand Bach.12 Their data are replotted in Fig. 2.19. The shape of the observed curve of hydrogenvs residual water content would seem to indicate a relationship similar to that predicted bySievert's law. In fact, a very close fit can be obtained through empirical calibration of thedilution term in equation (2-28). This, however, implies unreasonable amounts of CaCO3.Carbon monoxide in addition to that delivered by carbonates could be formed by oxidation ofcarbon. Again, an unreasonable amount of carbon loss would be required. Therefore, it mustbe concluded that further research is needed for a proper interpretation of the factors control-ling hydrogen pick-up in SA welding.

SAW (low-alloy steel)

Water content, wt%

Fig. 2.19. Hydrogen pick-up in SA welding at different water contents in the flux. Data from Evans andBach.12

Example (2.3)

Consider SA welding on a thick plate of low-alloy steel under the following conditions:

The flux contains 0.04 wt% H2O and is consumed at a rate of 0.6 g per g weld deposit.Estimate both 'potential' and 'equilibrium' hydrogen levels when the total oxidation loss ofcarbon in the weld pool is 0.03 wt%.

SolutionFirst we calculate the total amount of fused parent metal and weld deposit formed on welding.From equations (1-75) and (1-120), we have:

and

When the dilution ratio DI(B + D) is known, it is possible to calculate the total flux con-sumption per gram fused weld metal:

The 'potential' hydrogen level is thus:

If we assume that all CO produced by reactions between dissolved carbon and oxygen isinfiltrated in the arc column, the following balance is obtained:

Total number of moles of CO per g fused weld metal:

Total number of moles of H2O per g fused weld metal:

This gives:

Since the effective reaction temperature of hydrogen absorption in SA welding is not known,the maximum solubility of hydrogen at 1 atm total pressure is taken equal to 33 ppm, similar tothat in the previous example. By inserting this value in the expression for Sievert's law, weobtain:

In practice, the 'potential' hydrogen level represents an upper limit for the hydrogen con-centration which cannot be exceeded. Thus, the contradictory results obtained in the presentexample clearly illustrate the difficulties involved in estimating the effective partial pressureof hydrogen in SA welding.

2.8.4.4 Implications of Sievert's lawAn important implication of Sievert's law is that the fraction of hydrogen picked up from thearc atmosphere is very high at low hydrogen pressures:

(2-32)

As seen from equation (2-32), the first traces of hydrogen added to the atmosphere arecompletely absorbed in the metal. At increasing partial pressures the fraction of hydrogenpicked up in the metal will gradually decrease, finally attaining a threshold of (SH/2) in thecase of pure H2. This shows that the concept of 'potential' hydrogen content frequently used tocharacterise filler materials (see Fig. 2.11) is a dangerous one, since the rates of absorption areso different in the high and low ranges of the hydrogen potential.

2.8.4.5 Hydrogen in multi-run weldmentsSo far, no standardised method is available for the determination of hydrogen in multi-layerwelds. Early measurements by Roux,13 using an arrangement similar to that subsequentlyadopted in French standards, indicate a constant ratio of extracted hydrogen to the mass offused metal, regardless of the number of passes. If hydrogen is reported on the basis of depos-ited metal, this ratio may vary by a factor of 2.5 when comparing a deposit made in five passesto a single bead.

Exploratory measurements of local hydrogen contents in large-size joints have been madeby Skjolberg,14 who butt welded a 40 mm plate with a self-shielding flux cored wire at aninterpass temperature of 2000C. Samples were cut from a refrigerated part of the weldment atmid-thickness, including positions in the weld metal close to the fusion line and samples in theHAZ. His results are summarised in Table 2.3.

Normal testing of the filler wire according to ISO 3690 gave fused metal hydrogen contentsof 3.3 ppm (diffusible) and 1.7 ppm (residual). A comparison with Table 2.3 shows that themulti-run content of diffusible hydrogen is much lower than the corresponding ISO value,probably as a result of a high interpass temperature which facilitates loss of hydrogen to thesurroundings through diffusion.

Table 2.3 Measured hydrogen contents in multi-run FCA steel weldment. Data from Skjolberg.14

Condition

As-welded

PWHT*(4h/150°C)

Weld Metal

0.6 ppm diffusible0.9 ppm residual

0.35 ppm diffusible2.25 ppm residual

HAZ

Distance from fusion line (mm)

Oto 5

0.25 ppmdiffusible

0.15 ppmdiffusible

5 to 10

0.15 ppmdiffusible

0.15 ppmdiffusible

10 to 15

0.10 ppmdiffusible

*Post weld heat treated.

2.8.4.6 Hydrogen in non-ferrous weldmentsThe solubility of hydrogen in metals and alloys of industrial importance increases with tem-perature, and passes through a maximum in the vicinity of the boiling point, where the oppos-ing trends of increasing solubility and increasing dilution by metal vapour balance. Solubilitycurves for hydrogen in aluminium, copper, and nickel up to about 22000C have previouslybeen presented in Fig. 2.7.

Since all these metals can dissolve considerable amounts of hydrogen, the risk of hydrogenabsorption during welding is imminent if moisture is present in the shielding gas. Resultsobtained from arc melting experiments with Cu, Al, Ni in Ar-H2 gas atmospheres indicate thathydrogen is absorbed at a high temperature zone under the arc and is transported by fluid flowto the outer, cooler regions of the pool.15 Rejection of the gas in the supersaturated outerregions is slower than the absorption in the hot zone, so the gas content throughout the poolapproximates to that in the absorption zone. Typical estimates of the effective reaction tem-perature of hydrogen desorption (based on the Sievert's law) gave the following result:15

Copper: 16500C

Aluminium: 19000C

Nickel: 19000C

At present, it is not known whether these reaction temperatures also apply to conventionalGTA or GMA welding of the same materials or are mainly restricted to the operational condi-tions employed in the arc melting experiments.

2.9 Absorption of Nitrogen

It is generally recognised that interstitial nitrogen embrittles steel (e.g. see discussion in Chap-ter 7). In steel weld metals the associated loss of toughness due to free nitrogen has beenattributed to solid solution hardening and dislocation locking effects. In addition, excessivenitrogen pick-up can cause porosity in steel weldments because of gas evolution during solidi-fication.

2.9.7 Sources of nitrogen

Since the total nitrogen level in most welding consumables and shielding gases is quite low,the main source of nitrogen contamination is air infiltrated in the arc column. For this reason,the weld metal nitrogen content is very sensitive to variations in the operational conditions(e.g. arc length, electrode stick-out, shielding gas flow rate etc.). The overall reaction of nitro-gen absorption is similar to that of hydrogen:

(dissolved) (2-33)

(2-34)

By introducing the equilibrium constant K4 for the reaction, we get:

where SN is the maximum solubility of nitrogen at 1 atm total pressure,/^ is the activity coef-ficient, and pNl is the resulting partial pressure of diatomic nitrogen in the gas phase.

The solubility of nitrogen in liquid iron is approximately given by:

(2-35)

where T is the temperature in K.At 1600 and 20000C, this equation gives equilibrium concentrations of 446 and 465 ppm,

respectively. In alloyed steel containing large amounts of nitride-forming elements (e.g.austenitic stainless steel), the activity coefficient of nitrogen fN is about 1/4 and hence, thesolubility will be about 4 times higher than that calculated from equation (2-35).

From a primitive model of pseudo-equilibrium between gaseous N2 and dissolved N a maxi-mum solubility of about 465 ppm would be expected in welding under 1 atm total pressure.Thus, the maximum pick-up of nitrogen in deposition of bare wire in air would be of the orderof 465A/OT8 ppm or 416 ppm. If a tentative estimate of air infiltration in the arc column ismade at 1 vol% N2, the expected pick-up of nitrogen would be 465 VoToT or about 47 ppm.

A comparison with the data in Table 2.4 shows that the measured weld metal nitrogencontents are much higher than predicted from Sievert's law. This implies that the mechanismof nitrogen desorption is different from that of hydrogen.

2.9.2 Gas-shielded welding

Information on the factors controlling nitrogen pick-up may be obtained from the work ofKobayashi et a/.,16 who examined the GMA welding process in a systematic manner. Some oftheir results are shown in Fig. 2.20.

Figure 2.20(a), for low-alloy steel, reveals that the square root relationship is a fair approxi-mation only for welding in mixtures of N2 and H2 (curve No. 5). Mixtures of N2 + Ar (curveNo. 3), N2 + CO2 (curve No. 2) and N2 + O2 (curve No. 1) show increasing deviation from thepredicted behaviour. Pure N2 under reduced pressure gives a curve (No. 4) of an entirely differ-ent shape including a maximum at pN ~ 0.05.

Table 2.4 Summary of measured weld metal nitrogen contents. Data compiled by Christensen.4

Similar features are seen from Fig. 2.20(b) for welding of stainless steel. Again, the devia-tion becomes more pronounced as the oxidation potential of the gas mixture is increased in thesequence H2-Ar-CO2-O2. Moreover, a comparison with Fig. 2.20(a) reveals that the displace-ment of the nitrogen concentrations in the presence of chromium is larger than expected fromthe calculated reduction of the nitrogen activity coefficient.

The trends observed in Fig. 2.20 have been confirmed by O'Brien and Jordan17 who studiednitrogen pick-up during CO2-shielded welding of low-alloy steel. As can be seen from Fig.2.21 (a) their curves are similar to those of Kobayashi et al.16 for short circuiting metal transfer,while a mixed spray/globular transfer gives a sharp rise of nitrogen absorption up to pNi = 0.3followed by a constant or slightly decreasing concentration (Fig. 2.21(b)). Both patterns areclearly not in accordance with predictions based on Sievert's law (equation (2-34)).

An interpretation of the observed trends should be made with a view to absorption of hy-drogen, where the concept of pseudo-equilibrium has proved useful for a semiquantitativeprediction. In both cases the molecular species H2 and N2 are known to dissociate in the arccolumn (see Table 2.1), and would therefore dissolve in the metal to an extent far beyond thesolubility controlled by pH or PN . The excess of dissolved hydrogen is probably released asgas at weld pool temperatures. This will also be the case with nitrogen in the absence of oxy-gen, as shown previously in Fig. 2.20(a) and (b). However, under oxidising conditions thedesorption of gaseous nitrogen becomes suppressed by the presence of oxygen at the gas/metalinterface and hence, nitrogen is retained at a level which by far exceeds the solubility limit at1 atm total pressure of N2. This has been confirmed experimentally by Uda and Ohno18 in theirclassic work on surface active elements (i.e. oxygen, sulphur and selenium) in liquid steel. Asimilar phenomenon was quoted in Section 2.8.4.1 from the work of Matsuda et al9 even inthe case of hydrogen, where increased entrapment of hydrogen was observed in the presenceof oxygen (see Fig. 2.15).

It appears thus that excessive absorption of nitrogen (and in some cases also hydrogen)should be interpreted as a state of incomplete release of solute, as described previously inSections 2.6 and 2.7. As a consequence, Sievert's law cannot be used for an estimate of nitro-gen pick-up in steel welding, unless the weld metal oxygen content is extremely low.

2.9.3 Covered electrodes

The nitrogen content of SMA weld deposits is known to be sensitive to variations in the arc

Welding Method Material Nitrogen Content (ppm)

SMAW (basic electrodes) Low-alloy steel 60-180

Stainless steel 550-650

SMAW (rutile electrodes) Low-alloy steel 200-350

Stainless steel 600-750

SAW Low-alloy steel 40-140

FCAW Low-alloy steel 125-275

GMAW Low-alloy steel 50-200

Nitr

ogen

con

tent

, wt%

Nitr

ogen

con

tent

, wt%

Fig. 2.20. Nitrogen pick-up in GMA welding at different concentrations of N2 in the shielding gas;(a) Low-alloy steel, (b) Stainless steel. Data from Kobayashi et al.16

Vol% N2 in shielding gas

GMAW (stainless steel)

(b)

Vol% N2 in shielding gas

GMAW (low-alloy steel)

(a)

Nitr

ogen c

onte

nt, w

t%

Nitr

ogen

con

tent

, wt%

(a)

Low-alloy steel

Experiment

Vol% N2 in shielding gas

Low-alloy steel

(b)

Experiment Fig. 2.21. Nitrogen pick-up in GMA weld-ing at different concentrations of N2 in theshielding gas; (a) Short circuting metaltransfer, (b) Mixed and free flight metaltransfer. Data from O'Brien and Jordan.17

Vol% N2in shielding gas

Nitr

ogen

con

tent

, ppm

length (voltage) because of the risk of air infiltration in the arc column. This point is moreclearly illustrated in Fig. 2.22, which shows that the resulting weld metal nitrogen level mayvary significantly from one weld to another, depending on the operational conditions applied.Consequently, the use of long arcs in SMAW should be avoided in order to prevent excessivepick-up of nitrogen from the surrounding atmosphere.

2.9.4 Submerged arc welding

In submerged arc welding the risk of air infiltration in the arc column is less imminent, sincewelding is performed under the shield of a flux. Hence, in multipass welds the filler wire itselfwill be the main source of nitrogen (see Fig. 2.23), while in single pass weldments the baseplate nitrogen content is more important because of the high dilution involved. The latter pointis illustrated by the following numerical example.

Example (2.4)

Consider SA (single pass) welding on a thick plate of low-alloy steel under the followingconditions:

Based on the 'rule of mixtures', calculate the weld metal nitrogen content. Assume in thesecalculations that the nitrogen content of the base plate and the filler wire is 0.005 and 0.012wt%, respectively.

SMAW (low-alloy steel)4 and 5 mm basic covered electrodes

Fig. 2.22. Natural fluctuations in nitrogen pick-up during SMA welding due to variations in the arclength. Data from Morigaki et al.19

Welder No.

A B C D

Wel

d m

etal

nitr

ogen

con

tent

, ppm

loss

gain

SAW (multipass steel weldments)

Nitrogen content in electrode wire, ppm

Fig. 2.23. Nitrogen pick-up in SA welding at different levels of nitrogen in the electrode wire. Data fromBhadeshia et a/.20

SolutionFirst we calculate the total amount of fused parent metal and weld deposit formed on welding.From equations (1-75) and (1-120), we have:

and

The 'rule of mixtures' gives us the nominal weld metal nitrogen content, which is definedas:

The above calculations show that the nitrogen content of single pass SA steel welds is closeto that of the base plate because of the high dilution involved. This is in agreement with gen-eral experience.

Next Page