ELECTROSLAG WELDING: THE EFFECT OF SLAG COMPOSITION

ON MECHANICAL PROPERTIES

by

JAMES S. MITCHELL B.Sc, Queens University at Kingston, 1973

A THESIS SUBMITTED IN PARTIAL FULFILLMENT OF

THE REQUIREMENTS FOR THE DEGREE OF

MASTER OF APPLIED SCIENCE

In the Department

of

METALLURGY

We accept this thesis as conforming

to the required standard

THE UNIVERSITY OF BRITISH COLUMBIA

August 1977

© James Mitchell, 1977 In presenting this thesis in partial fulfilment of the requirements for an advanced degree at the University of British Columbia, I agree that the Library shall make it freely available for reference and study.

I further agree that permission for extensive copying of this thesis for scholarly purposes may be granted by the Head of my Department or by his representatives. It is understood that copying or publication of this thesis for financial gain shall not be allowed without my written permission.

Department of

The University of British Columbia 2075 Wesbrook Place Vancouver, Canada V6T 1W5

Date 7 ABSTRACT

Previous studies of the properties of electroslag weld metal have been done using electroslag remelted ingots made under welding conditions.

This procedure assumes the electrical and thermal regimes of these pro• cesses to be equivalent. To test this assumption an experimental program was devised in which the remelted metal of an ingot and weld made with each of three slag systems was analysed and the mechanical properties examined. The results show that each process imparted different properties to the remelted metal by alloy and inclusion modification. Consequently the above assumption was proved invalid. Special consideration was given to the effect of inclusion composition and overall distribution toward mechanical properties.

ii TABLE OF CONTENTS

Page

ABSTRACT . ii

TABLE OF CONTENTS . . iii

LIST OF FIGURES vi

LIST OF TABLES viii

LIST OF SYMBOLS ix

ACKNOWLEDGMENTS . . . x

Chapter

I INTRODUCTION 1

1.1.1 The ESW Process 1

I.1.1.1 Applications of ESW ..... 1

1.1.2 Properties of ESW . 4

1.1.3 General Properties of Welds ...... 6

1.1.3.1 Weld Metal 6

1.1.3.2 The HAZ . . 8

1.2.1 Slag Composition 9

1.2.1.1 Liquid Slag Chemistry 9

1.2.1.2 Slag Reactions 10

1.2.1.3 Slag Electrochemical Reactions 16

1.2.1.4 Comparison of ESR-ESW

Slag Requirements . . . 17

1.2.1.5 Slag Characteristics 19

1.2.1.6 Welding Considerations 21 iii Chapter Page

1.2.2 Inclusions and Mechanical Properties 22

1.2.2.1 Inclusions 26

1.2.2.2 Mechanical Properties 29

1.2.3 Gas Porosity 31

1.2.4 Heat Distribution and Structure 33

1.3.1 Summary 35

1.3.2 Statement of the Problem 37

1.3.3 Experimental 38

II EXPERIMENTAL AND RESULTS . 40

II. 1 Materials . 40

II.1.1 Steel 40

II. 1.2 Slag 40

11.1.2.1 CaF2 43

11.1.2.2 A1203 . 43

11.1.2.3 CaO 43

II. 1.2.4 Si02 43

II. 2 Apparatus 44

II. 3 Procedure 46

11.3.1 Chemical Analysis 46

11.3.2 Metallographic Analysis 50

11.3.3 Mechanical Testing ... 56

III DISCUSSION 71

III. l Slag Effects on Alloy Composition 71

III.2 Slag Effects on Inclusion Population 76

iv Chapter Paee

111.3 Inclusion Distribution 83

111.4 Mechanical Properties 86

IV CONCLUSIONS 92

BIBLIOGRAPHY 95

APPENDIX I 98

APPENDIX II 101

APPENDIX III • 102

v LIST OF FIGURES

Figure Page

1 Schematic of Electroslag Welding Process 2

2 Hydraulic Press Housing 3

3 Electroslag Production of a Rotor 3

4 Horizontal Cross-section of Melt Bw 7

5 Effect of Slag Basicity on Resulfurization Ratio . . 13

6 Slag Basicity vs Sulfur Content of ESR Ingots for Various Remelting Power Modes 15

7 Electrode Tip Current Density vs Fill Ratio of most Electroslag Processes 18

8 Relationship of Shelf Energy and Volume Fraction of Inclusions 23

9 Effect of Second Phase Volume Fraction on Total Strain to Failure 24

10 Effect of Inclusion Shape on Mechanical Anisotropy of an as Rolled, Low Carbon Steel 25

11 Free Energy of Formation of Some Metal Sulfides . . 27

12 Example of Centre-line Cracking 34

13 Macrostructure of an ESR Ingot . 36

14 Experimental ESW Apparatus Used 39

15 Published Specifications of Alloy Welten 80-C ... 41

16 Electrode Guide Assembly, Schematic 45

17 Schematic of ESW Configuration Before Starting, Slag Removed ...... 47

18 Sectioning of ESR Ingot and ESW Metal for Metallographic and Chemical Analysis 49

vi Figure Page

19 Inclusion Distribution of ESR Ingots 52

20 Inclusion Distribution of ESW Metal 53

21 Apparatus Used to Deep Etch Samples for S.E.M. Survey . . 55

22 Orientation of Impact Specimens Cut from ESR

Ingot 59

23 Orientation of Impact Specimens Cut from ESW ... 60

24 to 29 Impact Energy, Lateral Contraction and Fracture

Appearance Data 61 to 66

30 Instrumented Impact Oscillograph and Tracing ... 68

31 Anodic Polarization Curves for Pure Iron 75 32 Relationship of [Al] and [0] in Equilibrium with Alumina of Various Activities 78 33a X-Ray Energy Analysis of an Inclusion in Ci . . . . 89

33b X-Ray Energy Analysis of an Inclusion in Cw . . . . 90

vii LIST OF TABLES

Table Page

1 Example of Weld Metal Properties of Heavy Section Joints ..... 5

2 Commercial Fluxes used for Electroslag Processes 20

3 Chemical Analysis of Alloy Welten 80-C, as

Received 42

4 Remelting Conditions of Ingots and Welds ...... 48

5 Chemical Analysis of Ingot and Weld Metal 51

6 Inclusion Area Fraction, Spacing and Diameters . 54 7 Inclusions Observed on Deep Etched and Ductile

Fracture Surfaces 57

8 Observed Morphology of MnS Type Inclusions 58

9 Instrumented Impact Energies of CVN Specimens

Tested at 170°C 70

10 Electroactive Surface Areas . 73

11 Summary of Impact Test Data 87

12 Value of. n for Initial Slag. Compositions 101

viii PRINCIPAL SYMBOLS

A in metal phase

A in slag phase

A as a pure liquid

A as a pure solid

activity of specie A in slag

activity of specie A in metallic solution

free energy change

standard free energy change

free energy change at specified temperature

equilibrium constant

parts per million

temperature

weight percent

ix ACKNOWLEDGMENTS

I would like to sincerely thank my research supervisors

Dr. A. Mitchell and Dr. E.B. Hawbolt for their assistance and guidance throughout this study.

The financial assistance of a National Research Council Scholarship is gratefully acknowledged.

x CHAPTER I

Introduction

I.1.1. The Electroslag Welding Process

Electroslag welding (ESW) is the deposition of a filler metal supplied as an electrode through a liquid slag to effect fusion of metallic members, usually plates, see figure 1. Energy for the process is derived from electric resistance heating of the liquid slag. The current path is between the electrode and parent metallic plates.

In practice the process is limited to a few orientations and electrode configurations. Weld metal is laid in place by gravity, thus welding can only proceed in the vertical or near vertical direction. The electrode must be of sufficient cross section to accommodate large electric currents and supply weld metal to the joint. This factor and also the alignment in the gap, to prevent short circuiting are critical. These conditions set a minimum to the weld gap and material thicknesses that may be used.

However there is no maximum gap or thickness applied to the process.

Thus electroslag welding is primarily used in joining material of large cross section with one pass.

I.1.1.1. Applications of ESW

Electroslag welding has been used for the assembly of castings to produce large machine parts. Hydroelectric turbines, pump housings and press frames have been cast in sections and welded into final products (Figs. 2,3).

1 2

ELECTRODE

RUN - OUT BLOCK

PARENT PLATE

SLAG BATH

METAL POOL

WELD METAL

STARTING SUMP

FIG. I •• SCHEMATIC OF ELECTROSLAG WELDING PROCESS. FIG. 2 : HYDRAULIC PRESS HOUSING. REF.43.

FIG. 3= ELECTROSLAG PRODUCTION OF A ROTOR . REF. 54. 4

Also the fabrication of rotors has been accomplished by electroslag welding a series of electroslag refined ingots prior to forging, eliminating much of the waste encountered in conventional production practice.

As with all technology popular with industry, the reason for the adaptation of electroslag welding is economic. A weld of large cross section may be constructed with a one pass technique resulting in a saving of welding hours. The soft thermal profiles characteristic of the process reduces the need for preheating. These advantages, despite the necessity of a post heat treatment required in some cases, are the reason why the I process is being used. l'

1.1.2. Properties of ESW

Mechanical properties required of large section welds are different from those of thinner welds. Tensile strength and elongation specifications vary with weld metal thickness, see table 1, usually the greater the weld section, the lower are these requirements. Bend tests are also performed on welded structures. This is a simple estimate of weld performance and is confined to thinner weld assemblies. The results from this type of test are empirical, the outer fiber elongation and nature of cracking only are reported. Another test often specified for welds is the impact test.

This test requires that a minimum absorbed impact energy be obtained from areas within the welded joint at specified temperatures. Again although this test is empirical, some relations to structurejand processing have been correlated.

As a result of the solidification sequence it experiences, weld metal has a cast structure. Of this, the grain size and nature of non-metallic TABLE I : EXAMPLE OF WELD METAL PROPERTIES OF HEAVY SECTION

JOINTS. FROM REF. I .

ASTM A 242 or A44I ASTM A58I3

MATERIAL (in.) <75 .75-1.5 1.5-4 4-8 < 4 4-5 5 - 8 THICKNESS (mm) < 19 19-38 38-102 102- 203 < 102 102-127 127-203

MIN. lO'PSI 70 67 63 60 70 67 6 3 U.T.S. I0e Kg/m2 49 47 44 42 49 47 44

MIN. IOsPSI 50 46 42 40 50 46 42 Y. S. IOe Kg/m2 35 32 30 28 35 32 30

MIN. ELONGATION % 22 22 24' 24' 2 1 ' 21 ' 21 '

DEDUCT 0.5% FROM ELONGATION ( UP TO 3.0 % ) FOR EVERY

0.5 in.(l3mm) INCREASE IN THICKNESS ABOVE 3.5 in ( 89mm) . 6

inclusions have a major influence on mechanical properties. Since joint design and metal composition may alter the weld structure, they are chosen in such a way that the welded joint meets the required minimum specifications.

1.1.3. General Properties of Welds

Welding of steel is often performed to join materials that have been shaped by plastic deformation. The weld metal structure is then very different from that of the parent material (Fig. 4). The weld metal was liquid when laid in place and will have a cast structure. The thermal cycle imposed by this process onto the parent metal will result in a heat affected zone (HAZ) adjacent the weld metal. The structures of this zone are a function of the heat treatment properties of the parent steel and the amount of thermal input to the welding process.

1.1.3.1. Weld Metal

The sequence of events that occur in the weld pool after it has been

laid in place can be described by solidification theory. Heat is lost from

the pool to the joint and other heat sinks, causing solidification to

initiate from these areas. As grains grow' into the weld pool, segre•

gation of some constituents in solution occurs and these compounds are last

to freeze. The resulting structure of electroslag weld metal consists of col•

umnar grains with non-metallic inclusions situated in the interdendritic spaces.

Other inclusions may precipitate from solid solutions as the metal cools

leaving an even dispersion of fine particles throughout the melt. 7

FIG. 4= HORIZONTAL CROSS-SECTION OF MELT Bw 8

Properties of weld metal are related to the cast structure.

Characteristic of this is relatively high ductile to brittle transition

temperature indicating brittle fracture behavior.

1.1.3.2. The HAZ

The heat affected zone of a weld is not subject to the chemical pro•

perties of the slag as is the weld metal. However it is a product of the

thermal schedule of the welding process. Thus when properties of a welded

structure are determined this zone is also analysed.

The HAZ may be subdivided into two main regions. The area adjacent

the weld metal, which is exposed to temperatures ranging between the solidus

and A^, will contain structures of recrystallization and grain growth. In

electroslag welding the area near the weld is exposed to sufficient thermal

energy and time that some inclusions will dissolve and reprecipitate upon

cooling. Other inclusions, the more refractory variety, may spherodize or

remain unchanged. In either case, the result is a coarser grain size and

inclusion population.

For the alloy of this region the cooling cycle of the welding procedure

will lead to a microstructure which may be described by the continuous

cooling transformation temperature (CCT) behavior. The long thermal cycles

of electroslag welding result in shallow quenching gradients. When these

are superimposed onto the appropriate CCT diagram, the microstructure, in

theory, may be predicted directly. This structure, resulting from the soft

thermal profiles, and the altered inclusion morphologies combine and result

in a general derating of properties in this area. 9

The second region of the HAZ is that which is exposed to temperature below Al. The structure here reflects the effect of tempering nearest the warm annealed region to slight recovery further from the weld.

1.2.1. Slag Composition

Reactions between slag and metal have been the subject of many in• vestigations into steel refining practice. The influence of slag composition 2

on steel chemistry is well understood . Consequently the effect on steel

properties imparted by the slag via the alloy composition is very important.

For example, a slag may be designed to promote the reduction of oxygen into

the metal resulting in a large oxide or gas porosity fraction in the solidified

alloy. These constituents lead to reduced mechanical properties and are

described in the following sections.

Contrary to conventional steelmaking practice the interaction of slag,

metal and atmosphere for the electroslag processes is not well understood.

Many thermodynamic parameters such as temperature, hydrodynamic behavior and

actual slag composition are difficult to measure. Of these, slag composition

has a major effect on the liquid steel chemistry and therefore the metal

properties.

1.2.1.1. Liquid Slag Chemistry

Calcium fluoride, CaF2» is the basis of most slag systems used for the

electroslag process. It has the desirable properties of good electrical

conductivity, reasonable viscosity and while it will dissolve most oxides,

it is not a component of reduction-oxidation reactions at elevated tempera•

tures. Essentially, CaF9 behaves as an inert solvent into which slag 10

compounds as lime (CaO), silica (Si02) and alumina (Al^O^) will dissolve.

Once in solution the acid (oxide-ion acceptor) components, silica and 3 alumina, will undergo strong acid-base reactions to form stable ions . The

silica reactions,

2 2 5102 + 0" = Si03~ 1-1

to

-2 -2 -4

5103 +0 = Si04 1-2

and similar alumina reaction,

2 A1203 + 0~ = 2A102" 1-3

4

are believed to yield ions stable in basic slags . The existence of these

species in CaF2 melts has been verified by freezing point depression studies

on the appropriate systems"'. The source of oxygen to the above reactions

is the base type slag reactions discussed below.

Basic compounds (oxide-ion donors) dissolve in CaF2 slag by ionization. +2 -2

Lime in solution exists as the cation Ca and anion 0 . Thus when

describing the influence of each compound on the slag composition, or the

activity of these compounds in solution, it is important to consider the

actual ionic specie present in the melt. 1.2.1.2. Slag Reactions

For the major alloying elements in steel, thermochemical relations

are not sufficient to explain their mass transfer between electrode and

ingot . These elements include Cr, Si, Mn, Co, Cu and Ni and may be

present in quantities large enough to result in fast oxidation-reduction

reactions at the electrode/flux or metal pool/flux interface. In the present 11

work 60 Hz power was utilized, thus any direct current effects, such as gross concentrations gradients in the slag at the various slag interfaces, do not occur. However it is postulated that if the concentration of an oxidiz- able element in the metal (ingot or electrode) were large enough, the slight

60 Hz polarization effect would be sufficient to oxidize that element.

Thus a concentration gradient would develop at the anode. This occurs with

iron by the corrosion reaction

Fe = Fe+2 + 2e~ 1-4

The result of reaction 1-4 is a layer of slag saturated in (FeO) about the

anode^. Since the electrode/flux interface exhibits a smaller interfacial

area than the pool/flux interface, it develops a higher current density.

In effect the electrode/flux interface becomes the site of (FeO) saturation

during AC melting. An overall conclusion from the above phenomenon is

that oxidation at the anode can not take place if the anodic potential of

the reaction is greater than that of reaction 1-4. Also, if surface

depletion of an alloy occurs at the electrode it is by a mechanism of

anodic corrosion in preference to iron. Chromium has a lower polarizing

potential than iron and has been observed to develop a depletion gradient

at the electrode tip^. For other major alloying elements, like Si and Mn,

that have polarization potentials less than that of iron, the reaction

rate at the separate slag/metal interfaces is important. If the concen•

tration of the alloy element is high enough the rate of the appropriate

oxidation-reduction reaction may also be high. When the equilibrium co•

efficient is near to one, as it is for

(FeO) + [Mn] = (MnO) + [Fe] 1-5 12

then the difference in temperature between the electrode and metal pool will result in oxidation at the electrode and a nearly equal rate of reduction at the pool interface.

In practice, qualitative predictions of the effect of slag chemistry can be made using simple thermodynamics. The slag basicity ratio, or "V" ratio, applied to slag systems in steelmaking has proved useful in analysing

the effect of slag composition on metal chemistry. Early experiments have

concluded a sulfur distribution factor to be related to the acid and base g

constituents of a slag system (Fig. 5). This data indicates basic slags

lead to low sulfur contents in steel. The same work concluded CaF^ and FeO

contents of a slag and the slag temperature had little effect on the de-

sulfurization ratio. Also concluded was that MnO, MgO and CaO are equally 4

good desulfurizers. Reinterpreting these and other results, Turkdogan has

shown sulfur distribution between iron and basic slags may be written as [S] + (O"2) = (S~2) + [0] 1-6

In slags common to the electroslag process the dissolved oxygen in the slag

is related to the (CaO) content. The presence of lime in commercial CaF^ q

systems can not be avoided . Products of the desulfurization reaction,

(CaO) + [S] = (CaS) + [0] 1-7

(CaS) along with calcium (Ca) are highly soluble in CaF2 slags making them . * 10,11 excellent desulfurxzation systems

Alternately the existence of the less stable oxides like (FeO),

(SiO,,) and (MnO) in high proportions in the slag results in an increased oxygen

content of the metal pool10'"1"2. Simple equilibrium thermodynamics indicates

the reactions 13

FIG. 5 = EFFECT OF SLAG BASICITY v ON DESULFURIZATION RATIO. FROM REF. 4. 14

(FeO) = Fe + [0] 1-8

(Si02) = [Si] + 2[0] 1-9

(MnO) = [Mn] + [0] 1-10 are preferred in place of, for example, alumina reduction.

The influence of each component to the slag chemical reactivity has 13 14 been described as a function of the acidity ' . Slag compounds considered to be acidic are those which utilize the oxygen ion content and produce more complex metal ions. Reactions 1-1, 1-2 and 1-3 describe acid type reactions. Basic components are those which increase the oxygen ion activity in slags. These include the afore mentioned lime dissolution reaction CaO = (Ca+2) + (0~2) 1-11 and

MgO = (Mg+2) + (0-2) 1-12

The relative influence of these compounds has been empirically established.

In order of increasing acidity and relative strength the common acid

constituents are Fe20.j, 2 A^O^, 2 Si02 and 4 P20^. The numbers refer to

the moles of base necessary to neutralize one mole of the respective acid 2

compound . CaO and MgO are basic constituents of equal strength and CaF2 and FeO are considered neutral.

Slag basicity is only one factor affecting desulfurization, polarity

of the metal pool and electrode will also alter the sulfur removal mechanism^"* (Fig. 6). In practice electroslag welding is carried out in

either AC or DC mode, the latter being subdivided into DC electrode

negative (straight polarity), and vice-versa. Electrochemical aspects of 15

FIG. 6 SLAG BASICITY vs SULFUR CONTENT OF ESR INGOTS FOR VARIOUS REMELTING POWER MODES. FROM REF. 15. 16

DC remelting have been investigated and in general steels processed this way are found to be inferior to AC processed steel. In AC remelting the concentration polarization of ionic slag species at the electroactive surfaces is relatively small. This slight polarization is imposed by a net current rectification resulting from a saturated zone of (FeO) in the slag in contact with the electrode. For DC remelting polarization is at a maximum. The ensuing electrochemical reactions result In a large number of oxide inclusions in the remelted metal and induce the oxidation of many alloy components.

1.2.1.3. Slag Electrochemical Reactions 16

Careful experiments performed by Bell have led to the conclusion that simple chemical reactions between slag and metal cannot account for many of the compositional changes imparted to processed steels. Utilizing classical thermochemistry the computed equilibrium temperature of a slag species interacting with the metal is much higher than temperatures ex- 12 periences in the electroslag process . The reduction of alumina is an example of this. Thus electrochemical reactions were employed and some workers consider them responsible for as much as 75% of the overall chemical interaction.

Another condition supportive of the above observations is that chemical equilibrium during electroslag processing is not attained. The work of Fraser is evidence the rate controlling steps in this process is mass 6 transport . Thus in electroslag refining, a slower process than ESW, true equilibrium does not occur. Thus thermochemistry cannot account for the behavior of reactive species during processing, except in giving a quali• tative indication of the direction of a chemical reaction. 17

The overall effect of slag reactions results in an oxygen and sulfur content of the weld metal pool which can be related to the slag chemistry.

The levels of these components in the metal are however not calculable.

1.2.1.4. ESR-ESW Comparison

In view of the above chemical phenomenon a comparison of the electro• slag welding and refining techniques follows.

Refining is carried out in an electrically insulated mold, the current path being between the two active metal/slag surfaces, through the slag mass. Welding can be considered a live mold ESR process. The only difference is a part of the mold interacts electrically and chemically with the slag.

The current path is short and thus the thermal and electrochemical behavior of welding is very different from ESR. Also, in the refining process the pool/slag interface is an electroactive site. The bulk of the metal pool is available to interact with reaction products developed at this interface.

In electroslag welding the most electroactive site is adjacent the electrode.

This is cooler than the slag/pool interface of ESR ingots and lacks the fluid bulk as a source of reactants. Thus the pool/slag interface of the electroslag welding process is essentially the site of thermochemical reactions only.

Previously the effect of slag systems on the mechanical properties 3 of remelted steel has been investigated . However a problem arises when these results, derived in a small ESR mold using welding wire and high resistance slags, are correlated to weld metal performance. The implication that the electrode-ingot diameter ratio is the only relevant variable relating welding to refining cannot be supported with respect to the above argument (Fig. 7). The contribution of the electroactive mold toward 18

I.Oi T r——i r

AREA PROCESS REF. r. ESR 51 o ES R 52 .8 m BAR < ESW 47 ESR 53 ESW 52 ESW 47 WIRE I ESW 3

AY

_L 1.0 2.0 3.0 80 90 100 no CURRENT DENSITY (A/mm2)

FIG. 7 »'• ELECTRODE TIP CURRENT DENSITY vs FILL RATIO OF MOST ELECTROSLAG PROCESSES. 19

weld metal properties in ESW is significant and cannot be ignored.

1.2.1.5. Slag Characteristics

All slag systems must perform several necessary functions to the electroslag process, ESR or ESW. Slags used for welding are very different from those of refining operations, see table 2. In either process pro• perties of density, viscosity, surface tension, vapour pressure, specific heat and electrical conductivity influence the selection of an appropriate slag. The other slag requirement is chemical reactively which is imparted by the composition. Since current profiles in the slag bath are different between the welding and refining processes the slags used for each are different. The most pronounced difference being electrical conductivity.

To decrease slag conductivity welding slags contain a greater proportion of SiC^ and A^O^ and less CaF2 than slags common to ESR. The presence of silica and alumina increases the complex ion fraction of the melt and thus basicity decreases. These ions are less mobile than the smaller oxygen ion they consume. Since current conduction is a function of ionic transport in these slags the conductivity decreases. Also the decrease of basicity limits the refining action of welding slags, but generally this is not an important feature of selection.

Slag fluidity and stability are necessary to succesful electroslag welding and these can be adjusted to some extent with the CaF^, MgO, and

CaO contents. Also a necessary part of welding slags important to heat

transfer is the surface and interfacial tension existing between the slag/

atmosphere and slag/metal surfaces. Ideally these properties should be

low, as they are for acid type slags. Hence, MnO, FeO and again MgO and

Si0o are present in welding slag systems. 20

TABLE 2 : COMMERCIAL FLUXES USED FOR ELECTROSLAG PROCESSES

COMPOSITION ( WT. %) DESIGNATION USES

CoF, Al,05 CoO SiO; MgO TiOj MnO orMca

XI A A 60 26 10 4 EXPERIMENTAL

X 2 AA 60 17 20 3 ' ••

Y4 A 40 43 • 7 10 II

II — 75 5 10 5 5

— • 30 30 30 10 TOOL STEEL

— • 30 40 20 10 II

Y 2 A 30 34 26 10 II

T 1 C 50 34 4 12 II

— 80 5 10 5 II

— 55 25 15 2 3 DIE STEEL

Y 3 A 30 34 17 6 1 3 RENE 41

—- 35 "5 10 5 35 COLD START

AN 8 16 14 5 35 6 24 WELDING

AN 2 2 20 20 15 20 IS 10 » 5 AN 25 35 12 5 3 40 II Fe20 j

AN 10 20 20 1 0 2 0 30 —

FTs 7 5 3 3 47 17 25 WELDING 2 AN 8M 15 5 8 38 32 II N0;0

BV 8 42 26 24 8 II 21

There has been little attempt by industry to develop slags for each alloy that has been electroslag welded or remelted. However, some alloy categories and types of electroslag welding have been assigned various slag designations related to the composition of major flux components.

Only recently has one alloy, Rene 41, been allotted the slag Y3A"^, see table 2.

1.2.1.6. Welding Considerations

Electroslag welds are usually made with wire as the consumable electrode. To achieve the required mechanical properties from the weld metal region of the joint the alloy of the filler wire used is different from that of the parent material. A less frequently used type of electroslag welding used primarily for the manufacture of semi-finished products in• corporates a bar electrode often fabricated from the alloy of the parent stock used in the joint. Current density at the consumable bar electrode approaches the values encountered in ESR. This is unlike the high values observed with consumable wire ESW, see figure 7. Since electrode to plate gaps are small for bar electrode welding some features of welding

slags, i.e. low conductivity, may be required. Alternatively due to the

significantly lower electrode current densities, slag chemistries similar

to ESR systems may be used. The diversity and economic advantages of wire

fed ESW are lost when using bar electrodes but a wider range of welding

conditions and slag chemistries are possible with the latter. Another

consideration with wire welding practice is the commercial availability of

wire of the appropriate chemistry. This restricts precise matching of

mechanical properties throughout a welded joint. Also the electrical and 22

mechanical instability is a restriction. Large currents are passed through a ferrous alloy wire, typically 3 mm in diameter, into the slag phase.

Current density at the electrode tip is very high and slag design must allow for this. Problems with wire feed and electrical connection fix- turing have been met with specially designed machines, some of which are very complex.

1.2.2. Inclusions and Mechanical Properties

Many workers propose the majority inclusion population in ESR ingots are a result of the precipitation of non-metallics as the alloy solidif ies"^''

They maintain no inclusion present in the consumable electrode remains intact to join the ingot. The same arguments may be applied to the weld metal of an electroslag weld. As a result of slightly overheating the parent metal during welding some of this material melts and joins the pool without passing through the slag. Thus it is conceivable that some inclusion types, the more refractory variety, will enter the weld metal. In general, the inclusion fraction of the remelted metal of either process has been observed to be a fine dispersion of particles less than ten microns in size.

The mechanical behavior of steel is significantly affected by the

inclusion fraction. In general, inclusions have an adverse effect as shown

in figures 8-10. However not all inclusions relate the same effect to mechanical properties. In steelmaking practice a policy is made to minimize

the more harmful inclusion types as not all can be avoided. FIG. 8 ; EFFECT OF INCLUSION SHAPE ON MECHANICAL ANISOTROPY

OF AN AS ROLLED , LOW CARBON STEEL. FROM REF. 22. U> 24

EFFECT OF SECOND PHASE VOLUME FRACTION ON TOTAL STRAIN TO FRACTURE. FROM REF. 21 . 25

FIG. 10 » RELATIONSHIP OF SHELF ENERGY AND VOLUME FRACTION OF INCLUSIONS. FROM REF. 20. 26

1.2.2.1. Inclusions

Inclusions are the product of precipitation reactions from a liquid or solid solution. In section 1.2.1. the existence of inclusion forming elements in the melt is expounded. Using the theory of classical nu• cleation, as the dendritic solidification front advances into the metal pool it traps regions of stagnant liquid between the dendritic array. As freezing progresses, the concentration of segregation products increases in these regions. The high melting point inclusions, like alumina, will be precipitated first. With continued cooling and concentration of segre• gated species, the lower melting point oxides and sulfides will precipitate.

The later inclusion types will nucleate at existing inclusions resulting

in duplex type inclusions of which oxy-sulfides are an example. The last

form of Inclusion to evolve is a non-metallic precipitation from solid

solution. This type of inclusion, often sulfides, are extremely small and

evenly dispersed. The precipitation of iron sulfides is one case and occurs

along primary grain boundaries. In some circumstances this may form a 23

film like inclusion . This leads to low strength and ductility, and by

alloy design enough higher melting sulfide forming elements are added to

the alloy to precipitate in preference to iron from the melt. The standard

free energy of formation of some sulfides is shown in figure 11. The most

common element added for this is manganese. Recently more reactive sulfide

formers, Zr, La and Ce, have been introduced as these form very stable

solid sulfides in the melt. Since manganese is the principal sulfide

inclusion-forming addition to commercial steels it has been the subject of

many investigations. 0

-20h

-160" • • • — : ' ~—" 0 500 I'000 1500 TEMPERATURE (•C).

FIG. II FREE ENERGY OF FORMATION OF SOME METAL SULFIDES . FROM REF. 24. 28

Many workers have examined the formation and morphology of MnS in- 25-33 elusions in cast steels . A code for the various shapes of manganese sulfide particles observed was first proposed by Sims and has been retained throughout the literature. The morphological types are classed:

Type I - individual sulfides or oxy-sulfide particles, globular in shape and evenly dispersed throughout the matrix.

Type II - rod shaped sulfides, often clustered and described as eutectic. Type III - individual sulfides, angular (crystallographic) in appear- 26 ance. The effect of alloy elements common in steels and the solidification 29 rate on the formation of manganese sulfides has lead to a general con•

clusion that the shape of MnS inclusions is related to the solubility of 28 sulfur in liquid steel . Where the solubility becomes depressed, i.e. by the influence of solutes as carbon, silicon and aluminum, the sulfide type III will form as it precipitates at higher temperatures than type 25

I or II . All manganese sulfides form in the liquid occupying the inter•

stices of a dendritic solidification front. Slightly greater solubilities

of sulfur will render type II inclusions. The formation mechanism of these

is unclear. Sims concludes MnS forms as a continuous film over the primary grain boundaries which later breaks into rods as the structure cools. Other 26 30 33

workers ' ' describe type II formation as a monotectic reaction due to

the depressed freezing point imparted to the manganese sulfide by alloy

impurities. The origin of type I sulfides is generally agreed upon by all

workers. These have been observed as globular particles and the inter-

facial surface tension between the liquid manganese sulfide and iron is

high. They conclude the sulfide precipitates as a liquid from the segregated 29

liquid metal along the dendritic front. This precipitation reaction is aided by the existence of solid particles, usually oxide inclusions, onto which they nucleate from the segregate. The temperature of this reaction is below the liquidus of steel, therefore these sulfides contain impurities.

1.2.2.2. Mechanical Properties

The bulk of study on MnS inclusions has been toward their effect on the mechanical behavior of worked steels. Generally their contribution 21 22 2/j toward anisotropy of properties has been reported (see figure 9) ' '

In cast structures manganese sulfides have been observed at primary grain 25 26 boundaries and interdendritic spaces ' . The effect they Impart to mechanical properties is related to their overall distribution and shape.

For steel of a given inclusion volume fraction, evenly distributed equiaxed fine particles are favoured in spite of an increase in the absolute number. Large or clustered inclusion formations are detrimental to optimum mechanical properties as are non-equiaxed inclusions where isotropic be• havior is required. Thus manganese sulfide types I and III are desired

in place of type II. Although all sulfides are located in the area between primary grains of cast structures, type II occupies the largest area for a

given total sulfide volume. This creates a greater plane of weakness than types I or III and thusly type II is the most deleterious to mechanical 35 properties. Baker and Charles have observed interconnected colonies of

type II sulfides on the fracture surface of cast steel. This formation

provides a continuous crack path through the matrix resulting in inter-

granular fracture. This has only appeared in steels with relatively high

sulfur levels (about 0.28 wt. %). In all cases the situation becomes worse 30

when hot working leads to a further increase in the area to volume ratio and imparting greater anisotropy to mechanical behavior.

The overall effect of second phase particles within the metal matrix is to present a point of discontinuity to stress. For non-metallic in• clusions the inclusion matrix interfacial bond and thermal expansion coefficient relative to the matrix are important in determining the magni- 36 tude of this stress . Manganese sulfides have a large coefficient of thermal —6 —1 —6 —1 expansion (18 x 10 °C ), larger than steel (12.5 x 10 °C average) for 37 the temperature range less than 800°C . As a cast alloy cools the MnS inclusions will contract to a volume less than the cavity it occupies.

Since interfacial cohesion is low a void may form and so reduce tensile

stresses within the inclusion. This has been observed in the typically larger type I formations rendering these inclusions generally the second most detrimental to properties. The effect of the above inclusions and those with expansion coefficients less than steel (e.g. A^O^ and calcium aluminates) on the matrix has been 38

discussed in terms of fracture mechanics . The inclusion size, shape

and influence on the surrounding material was considered as a defect in

the matrix. The importance of tesselated stresses about inclusions of

smaller thermal expansion coefficients than steel is reported by Brooksbank 36 37

and Andrews ' as a mode of crack initiation. However they also report

inclusions with larger coefficients, like MnS, may also be considered defects

and impart a greater critical crack size to the applied stress. The

absolute values they calculate from plain strain fracture mechanics for

various materials is much larger than the size of inclusions found in

practice. Alternatively steel tending toward ductile rupture will fracture 31

by void formation, growth and coalescence. Manganese sulfides are usually

the larger species of inclusions in commercial steels, thus retaining the

greater volume fraction of the inclusion population. Since they exist in

a contracted state and have a low interfacial bond strength they are easily

separated from the matrix forming a void. This occurs at very small 39

strains . A subsequent increase in strain leads to void growth and

coalescence.

1.2.3. Gas Porosity

Porosity in an electroslag ingot is the result of the gaseous element

content of the liquid steel becoming insoluble during solidification and

precipitating as a bubble. The thermochemistry of gases in steel has been 2,14

dealt with in the literature and only the source of these gaseous

elements to the metal pool will be reviewed.

An advantage of the electroslag process is that the remelted steel is

never in direct contact with the atmosphere. Gaseous species picked up by 12

the metal must first be combined in the slag. Hawkins et al show high

nitrogen pick up by an ESR ingot at the beginning of a melt, but lower

(50 ppm) levels throughout the remainder of the ingot regardless of melting

mode. Nitrogen is soluble in calcium fluoride slags but its absorption

rate onto the metal pool surface is slow.

The hydrogen content of remelted metal is derived from three sources:

i hydrogen present in the electrode

ii hydrogenous compounds present in the slag prior to the start

of melting

iii the water vapour pressure of the gas phase in contact with the

slag 32

The hydrogen content of the electrode will be retained in ingots of 40 electroslag melts processed in contact with air . The high level of hydrogen pick-up at the start of an electroslag melt is attributed to the water content of hydrated slag compounds. The decrease in hydrogen level as an ingot is produced indicates the hydrogen from this source is con• sumed in the early stages of remelting. As the hydrogen content of the bulk of an ingot is constant a steady state condition is assumed. Thus a pseudo-equilibrium between hydrogenous species in the gas phase and metal pool becomes the predominant pick up mechanism. The slag composition directly affects the extent to which the second and third sources of hydrogen affect the remelted metal. In practice hydrated compounds of CaF^ and especially CaO are the principal sources of hydrogen. These compounds yield the very stable (OH ) ion in the slag. The mobility of this ion in slag has been assumed to control the rate of change of hydrogen concentration 40

in the pool . Overall, hydrogen is not considered too severe a problem to

ESW. The hydrogen contained in the weld metal has sufficient time to

diffuse away prior to the joint reaching room temperature.

Free oxygen levels in electroslag remelted metal must be kept low

to avoid the formation of blow holes. The presence of deoxidizing elements

in most alloys is sufficient to prevent this condition. However to obtain materials of high cleanliness, low oxide contents, the oxygen level of an

ingot must be reduced. When steel is melted through acid slags the for• mation of carbon monoxide blowholes is possible. The reaction

Si02 + 2[C] - [Si] + 2 CO 1-13

illustrates the importance of the interaction between [C], [Si] and oxygen.

Here dissolved silicon is the deoxidizing element. 33

The relationship between dissolved oxygen in the liquid metal and the 42

(FeO) content of the slag has been the subject of previous research

The ingot oxygen content increases as the slag (FeO) content. Electroslag processes exposed to air develop a higher (FeO) fraction than those shielded from air. Thus the atmosphere is a source of oxygen to the ingot. Alternate 16 oxygen sources include alumina equilibrium between slag and metal

(A1203) = 2[A1] + 3[0] 1-14

The effect of porosity on mechanical properties has been treated in the same way as inclusions. A hole, unlike an inclusion, will have an inter• facial bond strength and elastic modulus of zero. Inclusions like manganese sulfides are also treated this way (see section I.2.2.2.).

1.2.4. Heat distribution and Structure

The mechanical properties of an electroslag welded joint are greatly affected by the heat flow conditions under which it was produced. In con• strained electroslag welds solidification cracks result from tensile forces acting on the weld centre during fusion. In unconstrained welds cracking may result from a deep metal pool profile producing columnar grains oriented hori• zontally. This pool profile has been described as a "shape factor" and is a 43 function of slag composition and the welding schedule . Contraction of the weld metal just after solidification is greatest along the axis of columnar crystals. For horizontally oriented grains these forces are applied to the last region to freeze creating a centre line crack, see figure 12, and poor mechanical properties.

Heat flow in electroslag refining is very different from welding. The mechanical behavior of ingot metal reflects the orientation and size of the dendritic grains as compared to ESW metal. The remelted ingot metal is 34

FIG. 12= EXAMPLE OF CENTRE- LINE CRACKING FROM REF. 43. 35

subject to slower cooling profiles as the principal heat sink is the base.

Thus large mostly axially-oriented grains result and centre cavity formation does not exist (see figure 13).

Heat is generated in the electroslag process by an ohmic power drop through the slag. The magnitude of energy liberated in the slag is a function of slag conductivity which is a property of composition. Additions of alumina and silica to calcium fluoride decrease the conductivity and thus smaller currents are required to generate enough heat to melt steel. Lime additions make an insignificant change to the resistivity of these systems.

When relating properties of welds made with a variety of processes it is useful to compare the cooling profiles of the weld metal. ESR and

ESW have long metal cooling curves as compared to that of some arc weld metal deposits. The cooling gradients will determine the type of second phase segregation and primary grain size and are thus related to mechanical properties.

1.3.1. Summary

Although they are similar processes, electroslag welding and refining have some very different characteristics. Both processes involve remelting an electrode through liquid electrically conductive slag. The metal and

slag of the ESR melt Is held within a solid, thermally and electrically

insulating slag skin. On the other hand, the weld metal and slag of an

ESW are in direct contact with a part of the mold, the parent plates, and

interact with it. The influences these differences have toward the current

profiles of each process are very profound. This process difference leads FIG. 13= MACROSTRUCTURE OF AN ESR INGOT 37

to dissimilarities in the thermal and electrochemical behavior of each process, and consequently in the metal properties. 3

From the survey on mechanical properties the only investigator to compare electroslag Ingots to welds assumed the remelted metal of an ESR ingot made with welding-type slag was equivalent to weld metal of an ESW.

This implies that differences in electrical and thermal regimes impart the same influence to the remelted metal of both processes. Indeed, a system of ESW slag qualification has been proposed based on this assumption.

However, this is contrary to the conclusions of the review and should be resolved. From the survey of slag compositions we may conclude there is a significant difference between slags used for refining and welding. Thus an ingot and weld of each of the slag systems should be studied.

Remelted metal composition is related to slag composition for both processes. Minor element constituents of the alloy may be modified by interaction with the slag. Since mechanical properties are dependent upon the alloy element content, they too are altered by slag composition.

Of greater significance is the effect of slag composition on the inclusion character of the remelted metal. Consequently some ingot and weld metal properties should be analysed and the effect of slag composition be determined.

1.3.2. Statement of the Problem

From the foregoing discussion it is clear two important questions in the area of electroslag welding technology remain to be resolved. First, is it correct to assume that a weld and an ingot, melted through similar slags, will possess the same mechanical properties? If this statement is incorrect, 38

then one present standard method of slag qualification is in error. Second, may we assume that an electroslag weld has properties which are primarily controlled by the inclusion content? If so, then the slag composition has a direct influence on weld properties and may be optimized accordingly.

The program summarized below was undertaken in order to resolve these questions.

1.3.3. Experimental

An electroslag weld and ingot were made with each of three slag systems. The slag composition varied between that used for welding, one used for refining and a third system which was of neither extreme. The re• melted metal of each process was analysed chemically and metallographically and an assessment of mechanical properties was done. These tests and the results are described in the following sections.

All material for this study was obtained from a single plate of steel.

From this electrodes for both processes were cut and the parent plates for welding were obtained. The actual slag systems used were:

A. 80 wt 7, CaF2, 20 wt % Al^

B. 55 wt % CaF2, 35 wt % Al^, 10 wt % CaO

C. 55 wt % CaF2, 15 wt % Al^, 15 wt % CaO, 15 wt % Si02. FIG. 14- EXPERIMENTAL E.S.W. APPARATUS USED CHAPTER II

Experimental and Results

II.1. Materials

11.1.1. Steel

A single plate 1.5 inches (38 mm) thick of an alloy designated Welten

80-C produced by the Nippon Steel Corporation was used in all experiments.

This material was cut to make electrodes for welding and refining and the parent plate sections of the weld assemblies. Welten 80-C is a quench and tempered high strength low alloy structural steel with good weldability.

The general characteristics and chemical analysis of this alloy published by the producer are shown in figure 15 and table 3. The mechanical properties of this material are generally equivalent to ASTM A-514 and A-517. All cutting was done using a band saw to avoid any heating effects from flame cutting. Surfaces exposed to the electroslag process were ground free of oxide.

11.1.2. Slag

Slag compositions were made by mixing crushed and weighed samples of the various slag constituents. .After mixing the slag was stored in a drying oven at a temperature just above 100°C until required. Slags used in this program were based on calcium fluoride (CaF„), from Eldorado Nuclear, with

40 41

WEL-TEN 80C exhibits excellent resistance to stress cor• good properties make WEL-TEN 80C an ideal material rosion cracking because it does not contain nickel, an for semi-refrigeration pressure vessels for ethylene and element that contributes to H.S stress corrosion cracking. other chemicals. Its high notch toughness at low temperatures and other

P S Mo B Chemical Si Mn Cu Cr max max max max Composition max 0.18 0.15-0.35 0.60-1.20 0.030 0.030 0.15-0.50 0.70-1.30 0.60 0.006

Heat Treatment Quenched and Tempered

Available Thickness Range, mm 6 to 40, incl.

Yield Point, min kg/mm2 (psi) 70 (100,000)

Tensile Strength, kg/mm= (psi) 80-95 (113,800-135,100)

Thickness, mm ?i. min Test Specimen

6 to 13, excl. 16 JIS No. 5 Elongation Mechanical 13 to 21, excl. 22 )IS No. 5 Properties 21 to 40, incl. 16 JIS No. 4

Ratio of Bend Radius to Thickness, mm Specimen Thickness Bending Properties (180° bend radius) 6 to 32, incl. 1.5 Over 32 to 40, incl. 2.0

Charpy 2 mm V-Notch Thickness, mm Impact Value Notch Toughness Over 12 to 40, incl. 3.6kg-m (26 ft-lb) at -15 C

Maximum Weld Hardness Carbon Equivalent Thickness, mm Hv (to be tested when the value of carbon equivalent given at right 0.62 See Nole Over 12 450 max is exceeded) Nole. Carbon equivalent is calculated by the following formula:

C • "«Mn • "2)5i -'/(oNi-'-'/sCr '-'/(MO-'-'/HV

FIG. I5' PUBLISHED SPECIFICATIONS OF ALLOY WELTEN 80-C. TABLE 3 CHEMICAL ANALYSIS OF ALLOY WELTEN 80-C 1.5 inch ( 38mm ) THICK PLATE , AS RECEIVED .

SOURCE OF C Mn Si P S Ni Cr Cu Mo Al Nb V B OATA 0.18 0.6 0.15 0.03 0.03 0.70 0.15 0.60 .006 SPECIFICATION (max) 1.2 0.35 (mox) (mox) 1.30 0.50 (max) (max)

HEAT SHEET 0.12 .89 0.24 Oil .004 0.83 0.22 0 .33 0.04 .001

SPECTROGRAPHS 0.12 0.85 0.29 .014 .005 0.02 0.75 0.25 0.29 .076 .013 I DETERMINATION

ho 43

additions of alumina (A^O^) from Norton Company, lime (CaO) from Dynamit

Nobel and silica (S102) from General Electric. A triple beam balance was used to weigh the slag compounds with an accuracy within two percent.

11.1.2.1. Calcium Fluoride

Calcium fluoride was crushed and sieved to -6 +48 mesh, a size by experience found to be acceptable for the apparatus. Chemically, lime is the main impurity and can be found at a level of 500 ppm. This may vary between melts. Traces of silica and iron oxides have been detected but are of little consequence in slag systems containing dissolved oxides. The maximum impurity analysis of calcium fluoride, as received, is given below in wt. pet.:

Chlorine (Cl) .005

Sulphate (SO^) .01

Iron (Fe) ' .005

Lead (Pb) .005

Silica (Si02) ' -°5

11.1.2.2. Alumina

Electrofused alumina of 99.9% purity and a granular size of 16 mesh was used.

11.1.2.3. Calcium Oxide (Lime)

Recrystallized lime of 99.5% purity was used.

11.1.2.4. Silica

Clear fused quartz of 99.8% purity was obtained by crushing clean analytic tubing to a size of -6 +48 mesh. The main impurities are listed 44

below as an average in ppm:

5

2

50

CaO 7

MgO 2

4

4

II.2. Apparatus

9

The U.B.C. electroslag rig, as described by Etienne , was readily adapted to both ingot and weld production. A water cooled 3 inch (76 mm)

internal diameter copper mold was used to make an ESR ingot of approximately

10 inches (254 mm) in length. The mold and water jacket were electrically

insulated from the process resulting in a current path between the electrode

and base assemblies. The dimension of the electrodes used for refining was

1.5 x 0.75 inches (38 x 19 mm) in section and about 50 inches (1.27 mm) long.

For welding a fixture was installed which supported the parent plate

assembly such that the electrode bar and weld gap were aligned. The

electrode was positioned in the gap by a roll guide assembly that fastened

to the support fixture. This device (Fig. 16) was electrically insulated

from the weld assembly thus preventing a short circuit between the electrode

and plates. Water cooled, copper faced jackets were positioned on each side

of the gap to contain the slag bath. Electrical connection to the parent

plates resulted in a current path between the electrode and plates. 45

ELECTRODE

BOLT CLAMP IN A SLOT PERMITS ADJUSTMENT SPRING LOADED ROLLS JUL

RUN-OUT BLOCKS ELECTRICAL INSULATOR

7 h-WELD SUPPORT 8 ALIGNMENT JIG

FIG. 16 = ELECTRODE GUIDE ASSEMBLY , SCHEMATIC ( NOT TO SCALE ) . 46

Electrodes with a cross section of 1.0 x 0.75 inches (25 x 19 mm) and a length of approximately 50 inches (1.27 mm) were used to fill welding gaps of

I. 13 inches (29 mm) spread at the bottom which opened to 1.25 inches (32 mm) at the top. At the base of the assembly a starting sump made from the same alloy as the plates was welded into place (Fig. 17). Electrode bars for both processes were TIG welded during their fabrication to make electrodes of sufficient length.

II. 3. Procedure

The remelting program of ingots and welds was done using line frequency

(60 Hz) AC power. To start the process DC power was applied through a

compact placed between the electrode and sump or base plate. The cylindrical

compact measured 1.5 inches (38 mm) in diameter and about 1.5 inches (38 mm)

in height. It consisted of 60 grams of -325 mesh calcium fluoride mixed with

about 200 grams of steel turnings. When a molten slag pool was established

the power mode was switched to AC. All experiments were carried out in

contact with the atmosphere, no argon or other gas cover was used when

making either ingots or welds. Only material from the AC melt zone, ex•

cluding about 0.5 inches (12 mm) of the.top, was retained for analysis and

mechanical testing. The remelting conditions of the ESR and ESW experiments

are shown in table 4.

II.3.1. Chemical Analysis

The ingots and welds were sectioned as shown in figure 18. Samples

for chemical and metallurgical analysis were taken from the central region

of each melt. The chemical determinations were done using an ARL emission 47

RUN-OUT BLOCK

ELECTRODE

STARTING COMPACT

SUMP

(25) (32)

FIG. 17 SCHEMATIC OF ESW CONFIGURATION BEFORE STARTING, SLAG REMOVED. (NOT TO SCALE) TABLE 4 : REMELTING CONDITIONS OF INGOTS (i) a WELDS (w).

MELT * VOLT AMPS ELECTRODE FEED RATE INITIAL RATE OF SLAG SPEED MASS SLAG MASS ADDED DURING (V) (A) ( cm/ sec ) ( g/ sec) ( grams) MELT (g/min) x .01

25 1050 6.6 3.8 700 0 I A 1 B i 30 850 7.6 4.3 700 QO

27 1000 6.7 4.3 700 0 I C •

A w 24 1150 11.2 4.3 160 5

B w 24 IO5.0 9.2 3.5 160 5

4.2 160 C w [ 1050 11.0 5 I 24

* SLAGS « A - 80%CoF2, 20% Al203

B - 55 % CaF2 , 35 % Al203 , 10 % CaO

C - 55 % CaF2, 15 % Al20s , 15 % CaO , 15 % Si02 i - INGOT w - WELD 49

FIG. 18 * SECTIONING OF E S R INGOT AND ESW METAL FOR METALLOGRAPHIC AND CHEMICAL ANALYSIS. 50

spectograph of the E.S.Co. facility at Coquitlam, British Columbia. These specimens were about 1.0 inch (25 mm) square by 0.25 inches (6 mm) thick.

The surface analysed was first ground by the spectograph operator to comply with their experimental procedures. Alternate sulfur and oxygen contents were determined using Leco sulfur and oxygen analysing equipment. Table 5 lists these results along with the plate analysis.

II.3.2. Metallographic Analysis

Specimens for metallographic study were polished in the transverse plane, perpendicular to the axis of the ingot or weld. This surface was chosen as it best reveals interdendritic inclusion formations. A quantitative image analyser, Quantimit 720, was used to count and size inclusions over an area that totalled approximately one square millimetre. The survey was conducted at a magnification of 1220 X. Particles smaller than 0.5 microns in diameter could not be counted accurately and were excluded from the results.

Size distribution data are shown figures 19 and 20. The inclusion area fraction, average interparticle spacing and diameters are listed in table 6.

The same specimens used for the inclusion survey were deep etched in an apparatus shown schematically in figure 21. The etchant was 50 vol. % nitric acid in water and the bath temperature was not allowed to exceed 15°C.

Etching times varied between 8 and 10 minutes. After etching each specimen was quickly rinsed in a cold water bath then dried using alcohol. This procedure removed about 0.5 mm of metal from the surface and left the inclusions intact. Using a scanning electron microscope (ETEC Autoscan) the inclusions were observed and their constituent elements analysed with a multi-channel x-ray energy analyser (Ortec model 6200). The elements TABLE 5 : CHEMICAL ANALYSIS OF INGOT a WELD METAL

SPECTOGRAPHIC DETERMINATION " LE CO " uci T MtLI C Mn Si P S Ni Cr Cu Mo Al Nb S 0

.012 .001 40 A: .12 .83 .22 .014 <.00l .02 .81 .25 .30 .040

.005 119 A w .12 .87 .25 .014 .004 .02 .81 .25 .30 .047 .012

.80 .25 .30 .036 .012 .001 42 B i .11 .80 .18 .014 <.00l .02

103 B w .12 .84 .27 .015 .005 .02 .80 .25 .30 .086 .012 .008

34 .81 .24 .013 .001 .02 .80 .24 .29 .013 .Oil .002 j Ci I"12 .006 89 [ C w .12 .84 .29 .015 .004 .02 .80 .24 .30 .058 .Oil

.011 45 | PLATE .12 .85 .29 .014 .005 .02 .79 .25 .29 .076 .013

NOTE 5 All values are wt.-% ( except Oxygen analysis , which is ppm ) . C i

(>6 u= II. %)

j i u= 2 3 4 SIZE (M)

FIG. 19= INCLUSION DISTRIBUTION OF ES INGOTS. 50r A w B w C w

(>6JJ « 1.0 %.) (>6 » * 4.6 %) (>6 )X » 3.6 %) o 00 40 H ^ CD C 30 H O z 20 5

10

J L 1 J L T~1 i ' • ' 1 L 0 12 3 4 5 6 0 I 2 3 4 5 6 0 12 3 4 5 SIZE (p) SIZE (p) SIZE (p)

FIG. 20 : INCLUSION SIZE DISTRIBUTION OF ESW METAL TABLE 6 = INCLUSION AREA FRACTION , SPACING AND DIAMETERS .

INCLUSION «

MELT AREA AVERAGE DIAMETER FRACTION SPACING MEAN AVERAGE (%) (u) (u) (u)

A i .062 65 1.6

A w .166 44 1.4 2.0

B i .073 71 1.4 2.2

B w .174 64 2.1 3.0

Ci .094 89 1.5 3.1

Cw .242 60 2.1 3.3 55

SPECIMEN HOLDER L— THERMOMETER

COOLING SPECIMEN BATH

MAGNETIC r STIRRER

FIG. 21 = APPARATUS USED TO DEEP ETCH SAMPLES FOR S.E.M. SURVEY. 56

found to make up the inclusions in these samples are listed in table 7.

The above scanning electron microscope was used to examine inclusions on the ductile fracture surface of impact specimens. These specimens were charpy impact bars tested at 170°C. The results of this survey are also listed in table 7.

For these inclusion surveys, no element of atomic number less than 13

(aluminum) was detected by the x-ray energy analyser. Also, during the survey the type of the manganese sulfide inclusions were noted (see table 8), although sulfides generally were a small fraction of the inclusion population.

II.3.3. Mechanical Testing

Standard Charpy V notch impact test bars conforming to ASTM E-24 specifications were cut from the ingots and welds in the orientation shown in figures 22 and 23. For each melt the ductile to brittle transition temperature (DBTT) curves were determined. Each point represents a single test. From the broken impact bars the fracture appearance transition temperature (FATT) and the lateral contraction (LC) curves were derived

(Figures 24 to 29).

The fracture appearance data was obtained by making a tracing of the ductile regions of the fracture surfaces from enlarged photographs. Using a quantitative image analyser (Quantimet 720) the ductile area fractions were determined. An estimate of error using this method is less than ten

percent.

A travelling microscope was used to measure the contraction of shear

lips on impact bars. Accuracy of the measurement was greater than FATT

data. However since no alternate measuring of contraction could be used to

check this data, an estimate of error is difficult to assess. 57

TABLE 7 « INCLUSIONS OBSERVED ON DEEP ETCHED AND DUCTILE FRACTURE SURFACES.

DEEP ETCH FRACTURE MELT X-RAY PEAKS SIZE X-RAY PEAKS SIZE

Al 10-15 Al 6-8 A i Al.Zr 2-5 AI.Mn.S 7-12 AI.Zr.Mn.S 2-4 AI,Zr,Mn, S 3-5 Mn,S 3-4

Al 3-6 AI,Mn,S 5-6 A w AI.Mn.S 5-10 AI.Zr.Mn.S 4-6 Al.Zr , Mn , S 5-8 Al, Zr.Si, Mn.S 3-6 Mn.S 5-10 Mn.S 2-8 Mn.S.Si 5-6

Al 3-5 Al 4-8 B i Al.Zr 3-4 AI.Mn.S 7-10 AI.Zr.Mn.S 3-5 Al,Si,Mn.S 2-5 AI.Mn.S 2-3 Al.Zr.Si.Mn.S 2-5

Al 5-12 Al 2- 8 B w AI.Mn.S 2-8 AI.Mn.S 3-6 Mn, S 2-3 Mn.S 2-5

Al 7-13 Al 5-7 Al, Mn , S 3-5 AI.Mn.S 5-6 C i Mn, S 3-5 Si.Mn.S 1-2 Mn.S 2-5 Mn.S.Cr.Si 2-3

Al 2-5 AI.Si.Mn.S 3-5 Si.Mn.S 3-4 Cw AI.Mn.S 2-7 Mn, S 2-5 Mn.S 4-8 58

TABLE 8 : OBSERVED MORPHOLOGY OF MnS INCLUSIONS.

MELT MnS TYPE REMARK Ai I NOTE 1

Aw I PURE I NOTE 1 II PURE

B i I PURE I NOTE 1

B w I PURE H PURE

Ci I PURE I NOTE 2 rt PURE

Cw i PURE i NOTE 3 u PURE

NOTE I : SULFIDES WERE ASSOCIATED WITH PARTICLES CONTAINING Al AND Zr .

NOTE 2- SULFIDES WERE ASSOCIATED WITH PARTICLES CONTAINING Al .

NOTE 3= SULFIDES WERE ASSOCIATED WITH PARTICLES CONTAINING Al AND Si . 59

INGOT AXIS

FIG. 22 ' ORIENTATION OF IMPACT SPECIMENS

CUT FROM ESR INGOT. 60

WELD AXIS

FIG. 23 • ORIENTATION OF IMPACT SPECIMENS CUT FROM ESW. 61

FRACTURE APPEARANCE (%BRITTLE) o o o O O o cb

LATERAL CONTRACTION (mm)

u o

UJ cc

< or UJ a. ui

CVN ENERGY (ft.-lb.)

FIG. 24 = IMPACT ENERGY, LATERAL CONTRACTION and FRACTURE APPEARANCE for A i 62

FRACTURE APPEARANCE (%BRITTLE) o o o o o o CO (0 CM o

LATERAL CONTRACTION (mm)

a o

UJ cr < cr o_ UJ

CVN ENERGY (ft.-lb.)

FIG. 25 ' IMPACT ENERGY , LATERAL CONTRACTION

and FRACTURE APPEARANCE for Aw 63

FRACTURE APPEARANCE (%BRITTLE) o O O o o <\l o 00 10 -T-

LATERAL CONTRACTION (mm)

o o

or Z>

CVN ENERGY (ft.-lb.)

FIG. 26 « IMPACT ENERGY, LATERAL CONTRACTION

and FRACTURE APPEARANCE for Bi 64

FRACTURE APPEARANCE (%BRITTLE) O o o o o y oo ID v CM <- i i 1 i i 1 i i i 1 i LATERAL CONTRACTION (mm) O O in O m if)

CVN ENERGY

FIG. 27 •• IMPACT ENERGY, LATERAL CONTRACTION

and FRACTURE APPEARANCE for Bw 65

FRACTURE APPEARANCE (%BRITTLE) 2 o o o o _ S 00 «0 * (M ^ I I I I I II I I I *1 LATERAL CONTRACTION ( mm )

CVN ENERGY (ft.-lb.) FIG. 28 « IMPACT ENERGY , LATERAL CONTRACTION and FRACTURE APPEARANCE for Ci • 66

FRACTURE APPEARANCE (%BRITTLE)

o o o o o r O OO

CVN ENERGY (ft.-lb.)

FIG. 29 • IMPACT ENERGY , LATERAL CONTRACTION

ond FRACTURE APPEARANCE for Cw. 67

Impact specimens of each melt were fractured at a temperature well into the ductile region of the DBTT curve. An instrumented drop weight impact tester was used and the energy of fracture analysed. All specimens were tested at 170°C, using a silicon oil bath as the heating medium. An oscillograph of each test recorded the impact transducer voltage output

from semiconductor strain gauges as it varied with time during the fracture

event (Fig. 30). This was interpreted as striker, or tup, load versus

time from internal electronic calibration. Using a planimeter the area under the load-time trace was analysed. The areas representing crack

initiation and propagation were taken as the areas before and after the

maximum load point respectively. This and the following treatment were 44

procedures followed by Server for dynamic toughness evaluation of post

yield fracture.

A first approximation to the energy, Eo, represented by the area, A,

under the load-time curve may be found by Eo = A-v II-l

This applies to the areas corresponding to fracture initiation, propagation

or the total event. The striker velocity, v, at the time of impact with

the specimen is

v = (2.g-d)0,5 II-2

2 2

where g is the acceleration due to gravity (32.172 ft/sec , 9.806 m/sec )

and d is the mass drop height. Throughout this calculation the effects of

air resistance and frictional losses were assumed to be negligably small and

thus omitted. The value of Eo determined this way implies a constant striker

velocity during fracture. For brittle fracture, change in striker velocity 0.5 m sec. / div.

FIG. 30 = INSTRUMENTED IMPACT OSCILLOGRAPH AND TRACING . 69

is very small and thus any error to Eo will also be very small. However, where the absorbed energy is large, as it is during ductile rupture, a correction to Eo must be applied. The corrected energy, Ec, may be determined by

Ec = Eo(l-- Eo/4Es) II-3 where Es is the initial kinetic energy of the striker at impact and has the value

Es = m«g«d II-4

The value of m, the rest mass of the weight-striker assembly, was 3.14 lb (m) or 1.42 Kg. The corrected fracture energies are listed in table 9. 70

TABLE 9 ' INSTRUMENTED IMPACT ENERGIES OF CVN SPECIMENS TESTED AT 170 °C .

MELT NO. OF ENERGY ft.-lb. (N-m) TESTS INITIATION PROPAGATION TOTAL 124.0 Ai 5 30.6 (41.5) 93.4(126.6) (168.1) 88.7 A w 6 26.1 (35.4) 62.6 (84.9) (120.3)

108.2 Bi 4 29.1 (39.5) 79.1 (107.2) (146.7) 95.4 B w 5 26.8 (36.3) 68.6 ( 93.0) (129.3)

121.4 C i 4 30.1 (40.8) 91.3 (123.8) (164.6) 110.9 C w 5 28.8(39.0) 82.1 (III.3) (150.4) CHAPTER III

Discussion

Slag composition does not have a direct effect on mechanical properties.

These however are related by the effect slag composition has on the alloy and inclusion contents of the remelted metal, and the subsequent effect of these two parameters on mechanical properties.

III.l. Effects on Alloy Composition

Differences in alloy composition appear in table 5. Within the accuracy of the spectographic determination used the elements C, Cr, Cu, Mo and Nb are the same as the plate alloy content. Lower values of manganese and silicon were observed in all ingots whilst the weld metal contents of these elements remained relatively unchanged. A similar trend was observed with aluminum. Only the sulfur and oxygen analysis show the same difference as above, but no relationship between these and the plate contents can be established from the data.

The mechanism of alloy loss to the slag has been described in section 1.2.

Since a greater loss of the more oxidizable elements such as Mn and Si was observed in the ingots remelted through all slag systems, the slag com• position differences are not responsible for this effect. The magnitude of the difference in reactive element loss is different for the various slags

71 72

used. Since the absolute accuracy of the spectographic analysis technique is unknown, individual compositional differences between the ingot and weld of specific slag systems cannot be commented on.

Another mechanism that will explain the difference in alloy loss between the ESR and ESW process is postulated. Following the work of

Beynon^ on the electrochemical effects of the electroslag process, some very significant differences between welding and refining may be expected.

The electroslag refining technique used was an insulated mold process.

The effect of the difference in the metal-slag interfacial areas results in a greater current density at the electrode/slag than at the metal pool/slag

interface. The approximate interfacial areas of these surfaces are listed

in table 10. In electroslag welding the parent plates constitute a part

of the mold and are electrically active. The metal on this surface attains

at least the liquidus temperature of the alloy. As shown in table 10 the

differences of the interfacial areas, and thus current densities, are

greatest for electroslag welding. The polarization resulting from electro•

chemical reactions at these surfaces would develop a DC potential between

them and the electrode would become the anode. Since the magnitude of

this potential is related to the current density difference between the

electrode and metal sites it will be greater for electroslag welding.

As discussed in section 1.2. the polarization potential developed at

the electrode would not exceed that of iron oxidation,

Fe = Fe+2 + 2e~ III-l

For sufficiently high current densities the slag in contact with the electrode

tip would saturate with (FeO). This behavior was observed by Beynon in the

non-linear correlation between polarization potential and current density 73

TABLE 10 : ELECTROACTIVE SURFACE AREAS .

PROCESS ELECTRODE LIQUID METAL RATIO SURFACE

2 in2 (cm2) in (cm2)

ESR (i) 1.8 (11.4) 6.5 (41.9) 3.7

ESW (w) 0.8 (5.2) 5.6 (36.3) min. 7.0

6.0 (38.7) max. 7.5

* ESTIMATED FROM MELT Bw .

SCHEMATIC OF ELECTRODE TIP PROFILE .

ESR ESW 74

(Fig. 31). The result of the foregoing argument is that the slag adjacent to the electrode tip is high in oxidant, (FeO), for both electroslag refining and welding. Thus reactive alloy loss should appear in the data of both processes. As this was not observed the particular oxidation mechanisms of manganese and silicon shall be discussed.

Sufficient thermodynamic information exists that the equilibrium

constants of the Mn and Si deoxidation reactions may be calculated. The

proposed reactions are:

(FeO) + [Mn] = (MnO) + Fe III-2

2(FeO) + [Si] = (Si02) + 2 Fe III-3

= ^MnO) • >e] IIX_4

111 1 a(Fe0) • a[Mn]

111-3 a[Si] \"(FeO) /

The above equilibrium constants reduce to

KIII-2 = a(Mn°)/a[Mnl 111-6

K = a Si0 a Si 111-7 III-3 ( 2)/ t ^

if the slag at the electrode is considered saturated with (FeO). Using data

45 46 from Bodsworth and Kubaschewski the relationship between the equilibrium

constant and temperature may be calculated (see appendix AI).

log KII3;_2 = (6400/T) - 2.8 III-8

log Ki;[I_3 = (19200/T) - 6.8 III-9 75

FIG. 31 « ANODIC POLARIZATION {if) CURVES

FOR PURE IRON IN CaF2 + Al203 SLAGS . FROM REF. 7. 76

Both equilibrium constants decrease with increasing temperature. Fraser" has shown the temperature difference between the electrode/slag and pool/ slag interfaces of the ESR furnace will develop a situation where the rate of manganese reduction at the higher temperature pool interface is nearly equal to the oxidation rate at the electrode. There has been no precise measurement of temperatures during electroslag welding. However some

47 workers have observed the slag temperature during ESW to be much higher than the slag temperature of refining processes. This being the case, the above equilibrium constants are smaller for welding than refining. Thus more oxidation of manganese and silicon was expected during the refining

process than welding. This discussion relies on classical thermo•

dynamics and equilibrium conditions may not be achieved. The overall trend

however accounts for the difference in oxidative alloy loss between the

processes.

III.2. Effect on Inclusion Population

Other element recoveries were similar to that of manganese and silicon.

However the mechanisms of mass transport from the electrode to the metal

pool were different. . Aluminum and possibly zirconium were present in the

electrode primarily as stable oxides. The time and temperature, about 100°C

superheat, are insufficient to permit dissolution of these inclusions in

the electrode tip or droplet zones during remelting. Alternatively these

inclusions pass into the slag where they dissolve. Combining this mechanism

and the principles of oxidation from the preceding section, many workers"*"**'"^

conclude there are no inclusions In the remelted metal that originated in

the electrode. Inclusions detected in an ingot are the result of precipitation 77

reactions in the liquid or liquid-solid (mushy) region of the metal pool.

In electroslag welding it is conceivable that a second source of inclusions to the weld metal exists. Besides those formed by precipitation, some inclusions from the melted, or undercut, faces of the parent plates will enter the weld metal directly. The more refractory variety like alumina are in this category. The number of inclusions which enter the weld metal pool this way will be a function of the inclusion population and extent of undercutting of the parent plates. This accounts for some of the inclusion population observed in the weld metal. The majority, both oxides and sulfides, are a product of precipitation reactions during solidification.

The occurence of alumina inclusions in ESR ingots was studied by

Bell^. He concluded small evenly distributed alumina inclusions are a necessary constituent of ingots that had been remelted through a slag con• taining alumina. The calculated equilibrium diagram for the reaction

A1203 = 2[A1] + 3[0] 111-10

log KJH.JQ = (-64000/T) +20.5 III-ll is shown in figure 32. Each line represents the predicted equilibrium at

1700°C and 2000°C for various alumina activities in slags. For a given alumina activity the value of the equilibrium constant is proportional to the temperature. Since the slag temperature for welding is higher than refining, the metal pool of the welding process can be expected to contain a greater quantity of dissolved aluminum and oxygen. Table 5 shows this difference was observed in all slag systems.

A similar discussion may be used to describe the sulfur equilibrium between the metal pool and slag phase. Historically the mechanism of sulfur 78

FIG. 32 = RELATIONSHIP OF [Al] AND [o] IN

EQUILIBRIUM WITH ALUMINA OF VARIOUS

ACTIVITIES : FROM REF. 16 . 79

removal has been treated differently. For lime containing slags the 24 proposed desulfurization reaction is

(CaO) + [S] = (CaS) + [0] 111-12

AG° = +12.7 Kcal. (at 1600°C) 111-13 The equilibrium constant for this reaction is

a . *

In practice, sulfur removal has been analysed as a relationship between an empirically derived excess base quantity and the sulfur partition between

the slag and metal phases. The excess base quantities calculated for the

slags used in this study are negative (see appendix All). For these values

the sulfur partition, (S)/[S], would be close to unity. However some de•

sulfurization occurs, the extent of which cannot be determined by this type

of calculation.

The composition of the slag will change during the electroslag process.

Lime is present in all calcium fluoride based slags. In this study it is

conceivable that another source of lime was the formation of (CaO) from

slag-air reactions. Other compounds picked up by the slag during remelting

include (FeO). The source of oxygen to this slag constituent is also the

air. Thus the operating composition of the slag will be different from the

initial composition. The effect of these additions to a slag is to increase

'the slag basicity. The amount by which the basicity increases is in•

calculable as many essential variables, such as slag-air interaction and

anodic reactions, are unknown. For this reason the calculated excess base

quantity certainly does not relate to the desulfurization capacity of the

slag systems used. 80

The sulfur contents of table 5 show a general decrease as a result of both electroslag processes. This observation conforms to the above argument that all slag systems used in this study to some extent remove sulfur from the electrode stock.

Since both welding and refining was carried out exposed to air, sulfur 4 48 reactions between the slag and gas phase exist. Many workers ' have considered the sulfur distribution reaction

2 -2 h S2+ (0~ ) = h 02 + (S ) 111-15 at low oxygen ion activities, as in acid slag systems, this reaction favours the formation of gaseous sulfur. When slag-air systems are used the for- 4 mation of SO2 is most probable. Turkdogan defined a sulfide capacity of slag by the term Cs,

C = N (P p % s s ' 02/ S2) IH-16

where Ns is the atomic fraction of sulfide ions in the slag and ?Q2 and 7^ are the partial pressures of oxygen and sulfur respectively in the gas phase. From slag-metal and gas-metal equilibria he calculated

log Cg = - (2660/T) - 1.276 + v 111-17

The term v is a constant related to slag composition. By inspection, the

sulfide capacity of the slag increases as the temperature increases. Thus

the higher temperature slags of the welding process have a greater sulfide

capacity, and less sulfur is transferred to the air than in the ESR process.

In all cases the weld metal contained more sulfur than the ingot. This

cannot be due entirely to dilution of the remelted electrode metal by the

liquid steel from undercut areas of the parent plate. An estimate of the

fraction of weld metal made up by the melted parent plate was about 33 81

percent for the worst case, melt Bw (Appendix AIII). If dilution were the only cause of higher sulfur contents it would have to be about 70 percent to account for the difference between the weld and ingot analysis of the B experiment. Implied is the assumption that there is no sulfur refining of the parent plate metal by the slag occurs. Realistically some desulfurization would be expected requiring the dilution to be higher. 4

Following Turkdogan's analysis the above desulfurization reaction

(111-12) is essentially the oxygen-sulfur exchange reaction,

[S] + (0~2) - (S~2) + [0] 111-18

45

From Bodsworth the standard free energy of reaction is

AG° = 17,200 - 9.12 T cal./mole 111-19 and log KJJJ^Q - - (3750/T) + 1.996 111-20 -2

The reaction favours the products, (S ) + [0], as the temperature increases.

Thus the higher temperature welding process would result in greater sulfur removal from the metal if equilibrium were achieved. Since this was not observed, a desulfurization reaction more applicable to the elctroslag process must be used.

Fraser approached desulfurization of steel by combining the above oxygen-sulfur exchange reaction with the oxidation of iron Fe + [0] = (FeO) 111-21 +2 -2

The component (FeO) was considered to be ionized as (Fe ) and (0 ). The

resulting reaction is [S] + Fe = (S-2) + (Fe+2) 111-22 -2 +2 Treating (S ) as (CaS) and (Fe ) as (FeO) the equilibrium constant was 82

calculated as

(CaS) • a(FeO) K. 111-23 TII-22 a [s] also

K. = exp (2330/T - 4.07) 111-24 TII-22

This mechanism provides a decrease in desulfurization rate as the temperature increases. Equilibrium conditions are however implied which may not be the case and differences in the reaction constants at 1675°C and 2000°C are too small to account for the observed difference in sulfur removal.

A factor which does affect the above desulfurization reaction is the

(FeO) content of the slag. Because of polarization and the electrochemical behavior of slag-metal reactions (see section III-l), welding slags contain a higher (FeO) proportion than ESR slags. Since this source of (FeO) is

greater than that of the desulfurization reaction, the activity of (FeO) in

the slag is controlled by this mechanism. For both processes sulfur removal

is suppressed by this alternate source of (FeO). Since refining slags

contain less (FeO) than the corresponding ESW slags, more sulfur is

extracted from ESR processed steel than from weld metal. The resulting

sulfur content of the weld metal will be greater than the corresponding ingot

but less than the initial alloy.

The same difference in the slag (FeO) content between processes

accounts for the different metal oxygen contents. Added to this mechanism

are the reduction reactions of the more reactive alloy elements oxidized at

the electrode/slag interface (reactions III-2 and III-3). The amounts of 83

oxygen in the remelted metal that originate from these sources, and the aluminum equilibrium reaction (111-10), cannot be determined from these data. However, since the slags of the welding processes develop higher

contents of these oxidized alloy elements it is evident that the weld metal will reflect this with higher oxygen analysis as shown in table 5.

III.3. Inclusion Distribution

The effect which slag composition has on the remelted metal is also

observed by the inclusion character in the materials under study. Prior

to solidification the weld metal pool contains a greater concentration of

inclusion forming elements than the ESR metal pool. The weld metal contains

slightly more Mn and Si and much more Al, S, and 0. This observation

explains the difference between inclusion area fractions of table 6. For

all slag systems, the weld metal held a greater area fraction of inclusions

than the ingot metal. Thus features unique to electroslag welding, such

as higher operating temperatures and different slag electrochemical reactions,

result in a larger inclusion content.

The inclusion counts summarized in table 6 show a pattern which is only

related to the slag composition. These data indicate that inclusion area

fraction, size and possibly spacing, are related to slag basicity. The

excess base quantity of each slag is determined in appendix II. From these

calculations the slag system A (80% CaF2, 20% Al^O^) was the most basic and

C (55% CaF2, 15% Al^, 15% CaO, 15% Si02) the most acidic. The metal pro•

cessed with slag composition A, either ingot or weld, may be considered the

cleaner steel. This material had the least area fraction and size of

inclusions. The material remelted through slag system C had the greatest of 84

these two parameters. The inclusion distribution of slag B processed metal fell between A and C. Note the above trend is applied to ingot and weld metal considered separately. Also, an underlying assumption implied through• out this correlation is that slag compounds generated by each process have

an equal effect on the slag basicity of the individual melts. Specifically

that the (CaO) and (FeO) contents impart an identical influence to the

slag acid-base reactions for all systems. This assumption is reasonable.

The slag preparation and melting procedures were the same in all experiments.

Thus no difference in extra lime should develop between runs. Also the

(FeO) generated is related to the electrochemical characteristics dis•

cussed in section III-2, and was constant between electroslag welds and

ingots respectively. Thus as slag basicity increases, so does the general

cleanliness of the remelted steel.

The sulfur and oxygen analyses of table 5 slightly correlate to the 25 3

calculated slag excess base quantities. A conclusion of many researchers '

is that the type of manganese sulfide inclusion observed in steel is

related to the ratio of oxygen to sulfur analysis. This is also applicable

to these results. Remelted metal with [0]/[S] greater than 0.04 contained

exclusively type I MnS inclusions. Experiments Ai and Bi fell in this range.

In the remaining experiments the remelted steel contained both type I and

II inclusions and the [0]/[S] ratio ranged from 0.024 to 0.01. As this

ratio decreases further the morphology of MnS inclusions would be expected

to change to a population of only type II and eventually type III. For

practical purposes the appropriate values of [0] and [S] were not attained

and thus the higher sulfides not observed. Generally, the manganese sul•

fides were a small fraction of the total inclusion count. This was

irrespective of the absolute numbers obtained for the various experiments 85

and is a qualitative assessment as a segregated inclusion distribution analysis was not included in the procedure.

Unlike the MnS inclusion survey the composition of the remaining inclusions was not completely known. Major elements detected from these inclusions indicate their composition and variety differ between the processes and slag systems used. In all cases aluminum bearing inclusions were assumed to be alumina. Also inclusions that contained only aluminum

and zirconium, melts Ai and Bi, were assumed to be the oxides of these

elements. The composition of the remainder of the inclusions surveyed

could not be determined from the data. Some particles analysed indicated

a homogeneous composition while others show areas of local concentration

differences throughout the inclusion.

Table 7 lists the elements detected in inclusions observed on both

deep etched and fracture surfaces. Although most elemental combinations

found in inclusions are listed, some were less frequently found than others.

Inclusions that contained zirconium and in the one melt, Ci, chromium

were rare. Also inclusions greater than eight microns were rare, but

easily detected because of their size. Silicon bearing inclusions were

also rare in the A and B slag processed materials but accounted for a

large fraction of the inclusions observed in the C slag processed ingot and

weld metal. This was expected as the-C slag composition contained 15 weight

percent silica and thus the silica activity of this slag would be greater

than the others. The silica reduction reaction,

(SI02) = [Si] + 2[0] 111-25

obeys the same arguments applied to alumina reduction (section 1.2). Thus an

increase in slag silica activity is related to an increase in the dissolved 86

silicon in the metal pool and silica type inclusions would result from precipitation reactions during solidification.

III.4. Mechanical Properties

The data of figures 24 to 29 is summarized in table 11. There was much scatter among the transition temperature values and as such a correlation may not be drawn from these. The only consistent data is the upper shelf energy and a correlation between this parameter and the slag systems used will be discussed.

For every experiment the impact energies of the electroslag refined ingot metal was greater than that of the weld metal. The correspondingly lower inclusion area fraction of the ingot metal substantiates the trend shown in figure 10. Generally as the inclusion content decreases, the upper shelf impact energy increases. This correlation is true for both fracture initiation and propagation energy (see table 9). However the greatest difference between the weld and Ingot metal properties was in the propagation energies. In every experiment the change in propagation energy was greater than the corresponding change in initiation energy.

Thus the inclusion content of steel has a greater effect on crack propa• gation than on initiation.

Another trend apparent in table 11 was that the upper shelf energy increased as the acidity of the slag through which the steel was remelted increases. The impact energies of the C melt steel was greater than the

A melt steel, ingot and weld metal considered separately. This conflicts with the above conclusion that the upper shelf energy is adversely affected by the inclusion content of steel. The C slag processed material had the 87

TABLE II • SUMMARY OF IMPACT TEST DATA.

50% FATT 1 DBTT SHELF MELT L C C V N ENERGY CO (°C) (°C) (ft-lb.)

A i 88 94 98 121

Aw 98 95 98 86

B i 97 99 117 129

B w 86 8 8 109 94

C i 100 88 97 135

C w 103 100 117 1 12 88

highest inclusion content of the study. However in the literature a mechanism is proposed that accounts for this phenomenon and it has been applied to this data. 36 37

Brooksbank and Andrews ' conclude not all inclusions are detri• mental to mechanical properties in steel. They studied the effect of thermal expansion of inclusions on the tessellated stresses within the matrix in contact with an inclusion. A similar conclusion by Lyne and 49

Kosak was that resulfurized bearing steel contains relatively fewer

oxide type inclusions. These are the source of stress concentration sites

and are deleterious to fatigue properties^Also they conclude the

oxide inclusions encapsulated in sulfides are less effective stress

raisers than the single phase oxides. This is consistent with the findings

of Brooksbank and Andrews. They measured sulfide inclusions as having a

greater thermal expansion coefficient than steel. Thus the tessellated

stresses about an oxide inclusion surrounded by a sulfide would be reduced.

This argument can be applied to the present study. Although no inclusions

encapsulated in sulfides were found, their existence may not be discounted.

Some inclusions were observed to have a second phase component forming a

partial or thin coating about an inner phase (see Fig. 33). The composition

of these phases could not be separated with the technique employed.

Slightly higher sulfur contents and lower oxygen contents of the more

acidic slag processed ingot and weld metals would lead to the acceptance

of the Lyne and Kasak conclusion.

A less conjectural argument assumes the composition of inclusions

changes as the slag through which the steel was processed becomes more

acidic. The actual inclusion compositions could not be determined but the ALUMINUM SULFUR

CO

x LL. TOW v

FIG. 33 a = X-RAY ENERGY ANALYSIS OF AN INCLUSION IN Ci. (DEEP ETCHED

50 % HN03 , 1200 X ) FIG. 33 b X-RAY ENERGY ANALYSIS OF AN INCLUSION IN Cw. (DEEP ETCHED

50 % HN03 , 6000 X ) 91

observations of table 7 indicate the existence of elements such as silicon in the inclusions of melts B and C. From table 5 the oxygen contents of the ingot and weld metal, notably the weld metal, decreased as the slag acidity through which it was processed increased., This simultaneous decrease in oxygen levels and increase in the sulfur content and mechanical performance conforms to the conclusions of other workers^The lower oxygen levels may be directly interpreted as fewer oxide inclusions, whether they be aluminates or the less detrimental silicates. The higher sulfur levels impart a beneficial effect described in the previous paragraph.

Both sulfides and some silicates are known to have a greater inclusion to matrix cohesion than alumina and aluminate inclusions'^." For this reason

some improvement to fatigue properties has been reported for resulfurized

and acid processed steels.

The mechanism relating impact properties and inclusion character was

not the theme of this study. The general inverse relationship between

inclusion content and ductile impact energy has been reaffirmed. However

the detrimental effect of each type of inclusion and how they relate to

properties is important. As the inclusion composition is a function of the

slag composition through which the steel was remelted, then the slag

composition does effect mechanical properties. CHAPTER IV

Conclusion

Different alloy recoveries and inclusion contents were observed in the remelted metal of the welding and refining processes. The effect on impact properties was attributed to these parameters and the effect of inclusions on ductile fracture was reaffirmed. Since the only variable in this investigation was the welding and refining process the various process characteristics were studied, specifically the slag temperature and reactor geometry.

The slag temperature of welding was higher than refining although a precise value was not known. Relationships between temperature and equili brium constants provide a mechanism for greater manganese and silicon

recovery and aluminum and oxygen deposition in the weld metal than the

equivalent ingot. The reactor geometry influences the electrochemical

reactions at the liquid metal and electrode reaction sites. A greater

difference of current density characteristic to electroslag welding result

in higher polarization potentials between the active sites. In iron

systems the compound (FeO) is the product of this polarization and thus th

welding slags contain a higher fraction (FeO). This compound in the slag

suppresses the desulfurization reaction and contributes to an increase in

oxygen deposition.

These two process characteristics resulted in greater inclusion

contents and lower impact values for the weld metal.

92 93

The ductile impact energies varied as the inclusion content and type.

As the inclusion content increased for a given slag composition, the upper

shelf impact energy decreased. Instrumented impact testing showed the frac•

ture propagation energy to be most sensitive to inclusions. However as the

inclusion composition changed, notably as the sulfur and silicon levels in•

creased, higher impact energies were observed between welds or ingots

respectively. This reaffirms prior studies that the thermal expansion of

inclusions or the cohesive strength of the inclusion-matrix interface, or

both, affect mechanical properties. Thus chemical composition must be con•

sidered when relating inclusion content to properties.

The effect of slag on mechanical properties was considered in two ways.

Firstly the temperature and (FeO) content determine the extent of oxidation

at the electrode, introduction of elements, such as aluminum, which form

oxide inclusions, and the suppression of the desulfurization reaction.

Thus the inclusion volume fraction is related to these parameters. Secondly

the effect of slag acid-base behavior on slag/metal reactions was con•

sidered. For small changes in slag acidity little difference in alloy

recoveries was observed. However differences in inclusion composition were related to acidity and the particular slag composition. The presence

of silica in the slag and the resulting population of the less detrimental

silica type inclusions is an example of this.

From the results of the study we conclude that:

a. the mechanical properties of ingots and welds made using identical

slags and melting procedures are not the same. Hence an ESW slag cannot be

qualified by testing ingots remelted through it. The slag can only be

qualified by producing and testing electroslag welds. 94

b. the inclusion population of electroslag welds influences mechanical properties. However the relation is not a simple one and due account must be taken of the inclusion composition as well as the inclusion content. BIBLIOGRAPHY

1. American Welding Society "Welding Highway and Railway Bridges", AWS D2.0 - 69.

2. "The Physical Chemistry of Iron and Steel Making", Ward R.G., Edward Arnold Pub., 1962.

3. Zeke J., "Metallurgical Processes During Electroslag Melting Explained from the Viewpoint of Ion Theory" from Second Int. Symp. on ESR Technology, Carnegie-Mellon Inst., 1969.

4. Turkdogan E.T., J.I.S.I., JL79, 147, 1955.

5. Kojima H. Masson C.R., Can. J. Chem., 47^ 4221, 1969.

6. Fraser M.E. , Ph.D. Thesis, University of B.C., 1974.

7. Beynon G., Ph.D. Thesis, University of B.C., 1971.

8. Grant N.J., Chipman J., Trans. AIME, 167, 134, 1946.

9. EtienneM., Ph.D. Thesis, University of B.C., 1970.

10. Hawkins R.J., Davies M.W., J.I.S.I., 209, 226, 1971.

11. Rogers P.S., et al., "Metal-Molten Salt Solutions", in Physical Chemistry of Process Metallurgy, Pt. II, ed. St. Pierre G.R., Interscience Pub., 1959.

12. Hawkins R.J., et al., "Relevance of Laboratory Experiments to the Control of Composition in Production-Scale ESR", in Electroslag Refining, ISI Pub., p.21, 1973.

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17. Nafziger R.H., U.S. Bureau of Mines no. 669, p. 55, 1976.

95 96

18. Liddle J.F. , "Removal of Inclusions during ESR", in Chemical Metallurgy of Iron and Steel, ed. Bickle E.R. and Hawkins R.J., ISI, 1973.

19. Mitchell A., Ironmaking and Steelmaking, JL, (3), 172, 1974.

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52. Data from this work.

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t Appendix I

A.I Calculated Equilibrium Constants for Si and Mn

The equilibrium constants at 1600°C for the reactions

[Mn] + {FeO} = Fe + {MnO}

[Si] + 2{FeO} = 2Fe + {Si02}

61 62 have been calculated with the data from Bodsworth and Kubaschewski

The oxidation of manganese was first considered. For the reaction:

{Mn} + h02 = AI 1

the standard free energy change is

AG° = - 95,400 + 19.7 T AI 2 ref.62

To change solid MnO to a liquid product,

= {MnO} AI 3

the standard free energy change is expressed by:

AG° = 10,700 - 5.2 T AI 4 ref.61

The solution of manganese in liquid iron is: [Mn] = {Mn} AI 5

and: <

AG° = 9.11 T AI 6 ref.61

Combining the relations AI 1, AI 3 and AI 5 the resultant manganese oxidations

reaction is

98 99

[Mn] + h02 = {MnO} AI 7

The standard free energy of reaction is found by combining the equations AI 2,

AI 4 and AI 6

AG° = - 84,700 + 23.6 T AI 8

For the iron oxidation reaction

{FeO} = Fe + h02 AI 9

the standard free energy change is

AG° = 55,620 - 10.8 T AI 10 ref.62

Adding reactions AI 7 and AI 9 the oxidation of manganese by iron oxide may be

expressed as

{FeO} + [Mn] = [Fe] + {MnO} AI 11

and the free energy change is found by combining AI 8 and AI 10

AG" = - 29,080 + 12.8 T AI 12

The equilibrium constant represents the ratio of the reactants and products

as written and is expressed as

a fFe] ' {MnQ} AI 13

a „_T • a r[Mn] * 3{Fe0}

The equilibrium constant is related to the standard free energy of reaction

AG0 = - 2.303 RT log K

The value of R, the gas constant, is 1.98 cal'°C_1'mol 1. Thus the equili•

brium for manganese oxidation was related to temperature by 100

The same analysis was done for the oxidation of silicon

[Si] + 2{Fe0} = + 2[Fe] AI 14

AG° = - 87,960 + 31.1 T AI 15

The relationship between the silicon oxidation equilibrium constant and temperature is

19200 /- Q AT it log Kgi = —^ 6.8 AI 16

Due to inavailability of applicable data the reaction is written for solid silica as the product. In the slag systems used silica is very soluble and should be treated as liquid dissolved in a slag. These considerations would not alter the fact the equilibrium constant decreases as the reaction temperature increases. 101

Appendix II

A.II Excess Base Values of Slags

Excess base quantities were calculated for the initial slag com• position of each experiment. The relation between excess base and slag composition was taken from Bodsworth^, page 429.

E.B. = n CaO + n MgO + n MnO - 2n Si02 - 4n P205

- 2n A1203 - n ^&2°2

Here, n represents the number of moles of the oxide in 100 grams of slag.

The value of n for each oxide and the corresponding excess base quantity has been calculated and listed in table 12.

TABLE 12

VALUE OF n FOR INITIAL SLAG COMPOSITIONS

Si02 E.B. Slag A1203 CaO

A .196 - .39

B .343 .178 - .51

1 C .147 .267 .250 - .53 102

Appendix III

A.Ill Weld Metal Dilution

Dilution of weld metal in melt Bw:

From Figure 4 the area of weld metal was measured using a plani- meter. This figure is a photograph of a horizontal cross section from the upper end of weld Bw. Other welds, Aw and Cw, had less undercutting than

Bw and this analysis was not applied to them. From the location of figure 4 2 2 the area of the gap prior to welding is 1.5 x 1.25 = 1.88 in (12.1 cm ). 2 2

The measured cross section of the weld metal was 2.81 in (18.1 cm ).

Thus the dilution of the electrode metal by the plate is 100 (2.81 - 1.88)/2.81 = 33%