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METALLURGY 1. General Metallurgy
- Crystalline structure of metals and the allotropic states of iron
- Metal Alloys
- State Diagrams
- Solidification Structures
- Strain Hardening and Recrystallisation
- Behaviour of Metals at High and Low Temperatures
2. Mechanical Properties of Metals
- Strength
- Ductility
- Hardness
- Toughness
3. Steel Metallurgy
- Ferrous-Carbon State Diagram
- Structural Transformations in Steel
- TTT and CCT Curves
4. Manufacturing and Classification of Steels
- Methods of Steel Production, Refining, Casting and Lamination
- Classification of Steels
5. Heat Treatment for Steel
- Annealing
- Normalising
- Quenching
- Tempering
- Stress Relieving
WELDABILITY 1. Metallurgy of Welds in Steels
- The Weld Zone
- The Heat Affected Zone
2. Welding Defects
- Hydrogen in welding and cold cracking
- Hydrogen absorption in welding
- Hydrogen problems in welding
- Cracks in the HAZ (when cold)
- Hot Cracks
- Remedies for avoiding hot cracks
- Laminar Tearing
3. Heat Phenomena in Welds (Shrinkage, Internal Tensile Forces)
- Shrinkage in Welds
- Origin of residual tensions in welds
- Effects of shrinkage and residual tension
- Practical methods for attenuating tension
4. Welding Processes
- Electric Arc
- Welding processes using covered electrodes
- Submerged arc welding
- MIG and MAG welding processes
- TIG welding process
5. Weldability of materials
- Mild steels
- Low manganese alloy steels
- High resistance quenched and tempered steel
- Molybdenum-chromium steels
- Nickel steels
- Austenitic chromium-nickel stainless steels
- Austeno-ferritic stainless steels (DUPLEX)
- Aluminium and aluminium alloys
2
METALLURGY
1. GENERAL METALLURGY
Crystalline structure of metals and the allotropic states of iron
Crystalline structure
In their solid state, pure metals are made up of an aggregate of atoms that are alike, arranged in a
very specific order that is constant for each type of metal.
Unlike amorphous substances in which the elementary particles (molecules or atoms) are randomly
arranged in space, the spatial distribution of atoms in metals is governed by well-defined laws of
crystallographic symmetry – therefore metals have a crystalline structure.
The atoms that provide this specific structure are located at individual points in a fundamental
setting that is characteristic of a crystalline structure – the elementary cell.
In general, each cell can be considered as a polyhedron, the shape of which is determined by the
metal atoms positioned at its vertices.
Each elementary cell is surrounded by a certain number of identical, iso-oriented cells, with which it
shares a certain number of atoms that make up the links in a three-dimensional crystalline lattice.
Apart from a few exceptions, the lattices in pure industrial metals belong to the following three
crystallographic systems:
- cubic system with centred faces (Fig. 1)
- cubic system with a centred body (Fig. 2)
- compact hexagonal system (Fig. 3)
The cubic lattice with centred faces has cubic cells with an atom at each of the 8 corners of the
cube and an atom in the centre of each of the 6 faces.
The cubic lattice with a centred body has an atom at each of the 8 corners of the cube and an
atom in the centre of the cube.
The compact hexagonal lattice has cells that have a hexagonal prism shape, with an atom at
each of the 6 corners at the base of the prism, plus an atom in the centre of each base and three
atoms midway up the height of the prism at the corners of an equilateral triangle inside the
hexagonal prism.
Fig. 1
Fig. 2
3
Fig. 3
Allotropic states of iron
Allotropy is the result of the existence of different crystalline structure for the same chemical
species. Each of these exists in a stable state under specific pressure and temperature conditions.
Iron is a chemical element that can present itself in its solid state, in three stable allotropic forms at
different temperatures (Fig. 4).
Fig. 4
Alpha iron, which is stable from ordinary temperatures all the way up to 910°C, crystallises with a
centred body cubic system.
Alpha iron is magnetic at ordinary temperatures, but at 770°C it loses its magnetic properties. It is
then referred to as beta iron, but this should not be considered as a new allotropic phase as it
retains the same centred body cubic crystalline lattice as alpha Fe.
Gamma iron is stable between 910°C and 1440°C and is a real allotropic form of Fe as it has a
centred face cubic lattice. It is not magnetic and has a solvent power for C that is much higher
than alpha iron.
4
Delta iron is obtained due to allotropic transformation of gamma Fe when it is heated beyond
1440°C and is stable up to 1539°C, which is the melting point for pure iron. It has a centred body
cubic lattice, like that of alpha FE, but it is not magnetic.
Allotropic transformations, which, as we will see later, form the basis of the study of metallurgy,
come about due to the diffusion of Fe atoms and C atoms melted in the Fe. These atoms tend to
migrate in the material and therefore to homogenise its final structure, and this process is enhanced
by temperature.
We will also see what dangerous effects failure of these C and Fe atoms to diffuse during
solidification can have.
Metal Alloys
In practice, it is unlikely that pure metals will be used in steel construction. In order to obtain specific
mechanical, physical, and chemical properties, mixtures of two or metals are normally used, and
these are known as metal alloys.
Alloys can also include elements that are normally considered as metals, such as carbon,
phosphorous, silicon, etc.
There are two basic types of alloys:
- solid solutions
- juxtaposition alloys
Solid Solutions
Solid solutions can be obtained in two ways:
- by substitution
- by insertion
Substitution Solid Solutions (Fig. 5)
Substitution solid solutions are the solid solutions most commonly
found in metals, and occur if some of the atoms in the solvent
metal’s crystalline lattice are replaced by the same number of
atoms of the melted metal.
In general the solubility range is greater for metals with atomic
volumes that are very nearly the same, these tend to crystallise
into identical crystallographic systems, and are reasonably
close on the electro-chemical table of elements.
It has been shown experimentally that for the same crystalline
structure and vicinity on the electro-chemical table of
elements, substitution alloys are formed with wide ranges of
solubility where the metals have atomic diameters that do not
differ by more than 15% from one another.
Insertion Solid Solutions (Fig. 6)
Insertion solid solutions are formed by inserting one or more
atoms of the dissolved element into the links in the crystalline
lattice of the solvent metal.
In order for this to take place, the dissolved element must have
an atomic diameter that is significantly smaller (60% smaller)
than that of the solvent metal. The smaller the atom size of the
element to be dissolved (the solute) is in relation to the inter-
atomic spaces in the solvent, the easier it is for the solute to
insert itself into the crystalline lattice of the metal that
constitutes the basic matrix.
Only a few elements are able to satisfy this requirement when
dealing with iron – the most important are hydrogen, carbon,
nitrogen, and boron.
Fig. 5
Fig. 6
Solvent element
Soluted element
Solvent element
Soluted element
5
Experiments have also shown that central face cubic and
compact hexagonal crystalline structures are the structures
that are most favourable for forming insertion solutions.
Juxtaposition Alloys
Juxtaposition alloys are formed when the solubility
between the metals in their solid state is very limited
or non-existent.
In such cases alloys are made up of a number of
grains of the various metals or the possible partially
solid solutions in juxtaposition to one another.
These are known as juxtaposition alloys and include
the vast majority of industrial alloys.
One typical example is pearlite (Fe-C alloy, with 0,8%
C by weight). This is a juxtaposition alloy made up of
ferrite and cementite (Fig. 7).
Finally, it is worth noting that certain metals may
react with one another chemically to form
intermetallic compounds. This is the case with Fe-C
alloys, for example, that contain the Fe3C
(cementite) intermetallic compound, with 6,67% of C
by weight.
Fig. 7
State Diagrams
When a pure metal is cooled or heated, the corresponding temperature – time curve is not
constant but has two singular points between which the temperature remains constant during
cooling (Fig. 8).
These singular points act as pointers to the changes that take place at certain temperatures
(passage from the liquid to the solid state, from one allotropic state to another, that is, from one
crystalline shape to another, etc).
When looking at the cooling of a binary alloy (made up of two elements) we note that the
temperature curve first changes slope and then settles at a constant value for a certain time,
before beginning to descend regularly again. In this case we have three and not only two singular
points (Fig. 9). In addition, when the composition of the alloy is varied we note a resulting variation
in the position of these singular points. By drawing curves joining these points (Fig.10) on the basis of
the percentage concentration of the components in the alloy, temperature – concentration
diagrams are obtained the lines in which separate the fields of existence of the various alloy states,
and these are therefore referred to as STATE DIAGRAMS.
Fig. 8 Fig. 9
6
The state diagrams for binary alloys are drawn up by indicating the percentage concentration of
the two components on the abscissa (x-axis) and the temperature on the ordinate (y-axis). The two
pure metals will be at the two ends of the range and the characteristics for each intermediate alloy
will be identified by the intersection of the ordinate in relation to the alloy being studied, with the
diagram line.
The state diagram changes substantially in relation to the greater or lesser mixability capacity of
two components in the solid state. We will look at a case of binary alloys with total solubility in the
solid state and a case of zero solubility.
Fig. 10
Total solubility in the solid state
The diagram in Fig. 11 shows two singular points, one that marks the onset of solidification and one
at the end of solidification. The curves passing through these points are known as LIQUIDUS and
SOLIDUS, and they divide the diagram into fields of existence of the various phases.
Looking at an alloy with an M1 composition for example, during cooling the composition of the
liquid solution remains unchanged up to point P at temperature T1.
When the solution reaches this temperature, the first solid crystal separates and the alloy with the
highest melting point (Q first crystal) will have the highest proportion of the metal component.
Following this separation the composition of the liquid solution changes and becomes richer in
metal B.
As the temperature drops, the compositions of
the liquid and the solid follow the respective
curves so that, for example, at temperature T1’
the solid will have a composition Q’ and the
liquid a composition P’. When the temperature
drops further and crosses the “solidus” line at
point S, the last drop of liquid, with composition
R, solidifies and the composition of the solid will
be generally that relating to the M1 alloy
considered. At this point nothing else takes
place before ambient temperature is reached.
T
Fig. 11
solidus
liquidus
L+S
A B
P
P’
R
L
S
S
Q
Q’
T1
T’
T2
M1
7
Zero solubility in the solid state
In this case the diagram looks as shown in Fig.12.
Looking at an alloy M1 for example, involving two metals A and B, with a corresponding
composition of 25% of metal B, during cooling the alloy maintains its percentage composition until it
meets the “liquidus” line at point C and temperature T1. At this point the first grain of metal A will
form.
A reduction in temperature must result in a corresponding variation in the liquid solution. Since the
A crystals are continually being separated, the liquid solution is enriched with B and its composition
will follow the “liquidus” line, which means that, for example, at a temperature of T2 the
composition of the liquid phase will be D.
The composition of the liquid changes until point E is reached. At this point some B crystals will also
separate from the liquid and there will be three phases (A crystals – B crystals – liquid).
The temperature TE remains constant until the liquid phase disappears, after which the descent can
restart normally.
This means of solidification results in the alloy with zero solubility in its solid state, at ambient
temperature, being made up of metal A crystals which were the first to form and a mixture of small
juxtaposed A and B crystals, which were deposited while T was constant at point E.
This fine mixture of crystals (A + B) is known as EUTECTIC and point E as the EUTECTIC POINT.
If an alloy has a composition that is inferior to or superior to the eutectic point, it is referred to as
being a HYPOEUTECTIC ALLOY or a HYPEREUTECTIC ALLOY, respectively.
If it is considered as being a hypereutectic alloy, the phenomena that will take place during
cooling would be as indicated before, except that at ambient T the alloy would be made up of
metal B crystals and a mixture of small, juxtaposed B and A crystals.
T
Fig. 12
solidus
liquidus
L+B
A B
C
D
L
S
T1
TE
T2
M1
A+E
E
B+E (A+B)
L+A
8
Solidification Structures
In their liquid state, metal atoms have a high energy level given to them by the heat absorbed
during heating up from the solid state until the metal melts.
They are fairly free to move around in the heart of the liquid mass as they are obliged to maintain a
reciprocal distance but not their position, and they move at a speed that increases as the
temperature of the liquid increases.
When the liquid cools, the movement of the atoms decreases until, at a certain temperature and in
a generic point in the liquid mass, the inter-atomic forces of attraction gain the upper hand and
limit some atoms to a set position in relation to one another, thereby forming the first cell of the
crystalline lattice.
This is known as the first germ of solidification, and other cells form alongside it as the temperature
drops.
The growth of the crystalline structure does not take place casually but, as we have already seen,
according to well-defined crystallographic laws. In fact, the cells share the atoms in adjoining
faces and development takes place by branching out in the directions of the crystallographic axes
(solidification process said to be “arborescent” or “dendritic” - Fig. 13).
The result of the development of each germ of solidification is said to be “crystal”, “grain”, or
“dendrite”. Each grain contains a large number of elementary cells that are all oriented in the
same way in relation to the initial cell. This orientation varies from grain to grain.
Fig. 13
Oriented dendritic structures
When the temperature drops in a liquid mass on the verge of solidifying, this does not take place at
an equal rate in all directions as the specific external conditions result in directions with a greater
rate of change in temperature. The dendrites that form starting from the points that cool first
develop in the direction that has the steepest temperature gradient.
The grains therefore take on an elongated shape in a direction perpendicular to the isothermal
surfaces in the liquid in the process of solidifying.
In this case, dendritic crystallisation is referred to as being columnar, whereas when there are no
preferential directions of solidification it is referred to as being equiaxial.
In an ingot in its molten state for example, solidification starts in the area in contact with the cold
sides of the ingot mould. The perimeter grains are therefore elongated towards the centre,
providing a decidedly columnar structure.
In the centre, where cooling is practically unaffected by the cold sides, crystallisation is equiaxial.
In welding, due to the limited dimensions involved, solidification is completely dominated by a
columnar dendritic phenomenon and no space is available for equiaxial structures (Fig. 14).
9
Fig. 14
Dendritic segregation
When a pure metal solidifies, the material’s composition and structure remain identical for the
entire solidification time. This means that once crystallisation has been completed it is not possible
to distinguish which parts solidified first from those that solidified last in the crystalline structure; that
is, there is no segregation.
In practice however, perfectly pure metals are seldom used. The impurities and the alloying
elements cause phenomena that make it possible to differentiate between the parts that solidified
first and those that solidified last.
It should be noted, however, that when alloys solidify at a theoretical limiting velocity of zero, even
in the case of an impure metal or a metal alloy there are no phenomena of segregation. These
phenomena only occur at real solidification speeds.
In the case of an impure metal, the impurities generally dissolve in the liquid metal but not, or very
little, in the solid metal. This means that the first crystals deposited are pure from the remaining
liquid, which is enriched with impurities as solidification progresses.
Less pure metal is therefore deposited in the areas around the first dendritic grains and some of
these impurities will be segregated in the areas that solidify last – that is, around the grains.
Later we will see how this phenomenon can be of significant importance in welding as it is one of
the causes of a serious metallurgical defect (hot cracks).
10
Strain Hardening and Recrystallisation
Strain Hardening
Before introducing the concept of strain hardening, we need to introduce the theory of dislocation.
This theory is based on the supposition that a large number of elementary defects are to be found,
distributed randomly in the mass of metal materials.
A dislocation is an area in a lattice that is distorted locally where, over a radius of a few atomic
sizes, there are a string of atoms that are not opposite another string on the adjacent semi-plane
(Fig. 15 & 16).
Fig. 15: edge dislocation Fig. 16: screw dislocation
If there were no dislocation, any external shear force required to cause the two adjacent planes to
slide over one another would be greater, as all the atoms in the two planes would have to be
distanced from their position of equilibrium simultaneously. Since dislocation does occur, this
movement can take place gradually with the dislocation moving.
This explains why a metal’s tensile strength in practice is found to be lower than its theoretical
strength, which is specifically due to these dislocations.
Bear in mind, however, that everything that obstructs the movement of the dislocations tends to
increase the strength of the material.
Strain hardening is a fundamental characteristic in the plastic deformation of metal materials and
involves the continual increase in the force required to cause this sliding, as deformation increases.
This phenomenon can be seen in the “load – elongation” graph that is obtained in a mono-axial
tensile test and shows an ever upward trend on the curve over the plastic deformation stage.
Strain hardening is explained as the
interaction between the dislocations and the
increase in the number of these as plastic
deformation proceeds, for which gradually
increasing force is required to cause this
movement. Strain hardening that is obtained
by any cold plastic operation is very
dangerous as, in addition to resulting in an
increase in the ultimate tensile strength and
yield point, it also causes hardness to increase
and ductility and toughness to decrease (Fig.
17).
Fig. 17
11
Recrystallisation
This is the process that sees the crystalline lattice of material that has undergone plastic
deformation rebalanced as a function of time or temperature.
The name is derived from the fact that the reordering of the lattice starts from the germs of
crystallisation and their subsequent growth. Recrystallisation tends to be quicker as the cold
deformation increases.
The speed with which recrystallisation occurs is very slow at low temperatures. The minimum
temperature at which this takes place within a sufficiently short period of time is called the
recrystallisation temperature. Crystallisation tends to take place more quickly at higher
temperatures.
When plastic deformation of a metal piece is carried out at a temperature exceeding this
recrystallisation temperature, recrystallisation takes place immediately after plastic working (hot
working), whereas if the temperature is lower, plastic working causes strain hardening (cold
working) and recrystallisation does not occur other than after subsequent heat treatment.
Recrystallisation after cold plastic deformation therefore removes the effects of strain hardening,
lowering the strength and hardness of the material and increasing its ductility and toughness. As a
result of the new crystalline structure formed, this process also allows the grain of the metal to be
made finer.
12
Behaviour of Metals at High and Low Temperatures
High Temperatures
As temperature increases the material is able to relax and recrystallise, due to the greater speed of
diffusion and mobility of the atoms caused by the thermal energy. As a result, the material’s
capacity to harden reduces and so deformation takes place continuously, with the only obstacle
being the cohesion of the crystalline planes.
At sufficiently high temperatures, under the effect of a load applied even below the material’s
yield point, plastic deformation may occur and continue, even without any further increase in the
load. This phenomenon, which results in the material continuing to elongate until it breaks, without
any increase in load, is known as hot creep.
The creep phenomenon is greatly influenced by the test temperature, the degree of stress, and the
time for which the load is applied.
When the applied load is kept constant and the temperature is varied, the elongation – time graph
for a given material looks as shown in Fig. 18. The changing behaviour at the various temperatures
is due to the ratio between the speed of strain hardening and the speed of recrystallisation. In fact,
Fig. 18 shows three curves:
- one curve (no. 1) in which the increase in strength due to strain hardening the material is
dominant over the recrystallisation phenomenon
- one curve (no. 2) which, after the initial stretch, takes on a constant incline on the time axis and
shows a balance between the speed of strain hardening and that of recrystallisation
- one curve (no. 3) in which the second section, i.e. the part with the constant incline, is limited,
followed by a rapid upward turn in the curve, i.e. by a rapid increase in the speed of
elongation, which is always followed by breaking. In this case creep is characterised by
continual and almost complete recrystallisation.
One metallurgical factor that has a great influence on hot creep is grain size. It has been found
that a material’s resistance to creep in large grain metals is higher than that for those with fine
grains, when all other factors are equal, as the edges of the grains are involved in this process. In
fact, at high temperatures the edges of the grains facilitate the creep phenomenon both due to
their lower atomic density and because they constitute areas of attraction for dislocations that are
pulled apart, and lose their function of acting as brakes on the sliding between the atomic planes.
Thus large grain materials that have smaller edges provide greater resistance to creep.
Cr-Mo steels are very often used for high temperature settings and we will deal with these in the
section dealing with weldability.
Fig. 18
ε
t
T1
T2
T3
T3>T2>T1
Breakdown
13
Low Temperatures
With the onset of modern technology there is no lack of cases of metal structures such as
equipment or pressure vessels working at very low temperatures, even at –253°C (boiling point for
hydrogen).
The most noticeable changes occur in the mechanical characteristics of metals at low
temperature. These generally take the form of an increase in strength and a reduction in ductility
and toughness that varies in intensity (depending on the temperature), and appears to be due to
a greater difficulty in obtaining plastic creep (Fig. 29).
Whether or not a metal can be suitable for low working temperature applications is normally
determined in terms of its toughness at that temperature.
Experiments have shown that the factors that have the greatest influence on toughness are:
- the crystalline structure – metals with a centred body cubic system or compact hexagonal
system behave badly, while those with a centred face cubic crystallisation system behave well
(Fig. 31a/b).
- the manufacturing conditions and heat treatment, and therefore the fineness of the grain;
carbon-manganese steels with a fine grain can be used to temperatures of about –50°C, steels
with 3,5% Ni to about –90°C, and steels with 9% Ni to about –196°C.
14
2. Mechanical Properties of Metals
The mechanical properties of metals are taken as indicating all the properties that solid substances
(in this case metals) are found to have when a force is applied to them and that therefore indicate
the possibility of using a metal for structural purposes.
Metals are highly regarded in the construction industry as they are strong, hard, ductile and tough.
The combination of these properties can also be varied within extensive limits by adding alloying
elements, heat treatment, or mechanical action.
Strength
In the case of metal materials the term strength is used to mean a wide range of things, as it is used
in relation to the other characteristics of the materials (plasticity, hardness, toughness) and
therefore indicates the combination of all that is known about the relationship between applied
loads, internal forces (tension) and deformation.
Of the mechanical tests used, the single-axis tensile test is one of the best defined, in terms of
concept, to determine this relationship, as at least in theory it provides complete uniformity in the
application of the load over a rather well-defined area of material, known as the gauge length L0
of the test piece (Fig. 19).
Fig. 19
In order for the load to be applied uniformly over the entire gauge length, this must have a
constant section and the load must be applied at its ends in two areas known as heads,
connected to the gauge length very gradually.
The test piece subjected to a mono-axial load can have a circular, square or rectangular section,
or it may be in the form of a length of pipe when carrying out tensile testing of tubular products.
In general, tensile tests on a metal alloy include four stages or periods (Fig. 20):
- ultimate tensile strength (1)
- yield strength (2)
- ultimate load (3)
- period of elastic elongation
- period of yielding
- period of uniform plastic elongation (4)
- period of non-uniform plastic elongation or “necking down” (5)
If the load on the test piece is removed during elastic elongation, no permanent deformation
occurs.
The yield period marks the transition from elastic to plastic behaviour. The yield load (indicated as
Rs) is the load at which, for the first time, an increase in elongation occurs without a simultaneous
increase in loading, or with a decrease in loading. The period of plastic deformation occurs
15
immediately after yielding and ends when the maximum load or “ultimate tensile strength” is
reached.
The elongation of the test piece is localised over a short length of the test piece, where transverse
deformation (“necking down”) occurs and progresses until the test piece breaks (Fig. 21).
Fig. 20
Fig. 21: “necking down” in the test piece
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Ductility
Ductility is normally defined as a material’s capacity to deform plastically without breaking, and is
indicated by the extent to which a test piece subjected to a tensile test deforms plastically until it
breaks.
Good ductility is very important in a metallic material as:
- it is an indication of the material’s capacity to withstand cold deformation
- it gives an indication of the plasticity reserve that is available to withstand breaking due to
sudden overloading
- it means that you do not have to worry too much about the residual tension that is an
inevitable consequence of stresses induced by welding
- it gives an indication of the material’s strength in relation to lamellar tearing.
In order to test the ductility of the material, a test piece with a rectangular section is submitted to a
bending test (Fig. 22). Good results from this test indicate that the material is sufficiently plastic as
the outer layers of material have been elongated.
16
The size of the grain is very important in terms of ductility. As the grain size decreases, the material’s
ductility increases.
Fig. 22
Hardness
Hardness can be defined as a metal
material’s capacity to withstand a body
(indenter) seeking to penetrate the metal
material by deforming it plastically (by
compression).
The hardness of a test piece is determined on
the basis of the size of an indent left by the
indenter to which a specific force is applied.
There are various methods for measuring
hardness that differ in terms of:
- the shape of the indenter;
- the extent of the force applied;
- the method used to evaluate the
hardness.
Fig. 23 - durometer
Brinell Hardness
17
A load “F” is applied to a ball made of
hardened steel, and is kept in place for a
certain time. Once this time has expired, the
load is removed and the maximum diameter
of the circular imprint the sphere has left in
the material is measured.
The Brinell number is given by the ratio F/S
between the load and the area of the
imprint, and is indicated by the symbol HB
(Fig. 24). The Brinell Test is not reliable for
materials with an HB value that exceeds 450.
Fig. 24
Vickers Hardness
A load “F” is applied to a diamond indenter
having a pyramidal shape with a square base
and angle at the vertex of 136°, with the load
being applied for a certain amount of time.
The load “F” is removed and an indentation is
found that is considered as being a square-
based pyramid with the same vertex angle as
the indenter.
The area of this indentation is then calculated
after having measured the diagonal of the
base of the pyramid. The Vickers Hardness is
the ratio F/S between the load and the area
of the imprint, and is indicated by the symbol
HV (Fig. 25).
Fig. 25
Rockwell B and C Hardness
An initial load F0 is applied to an indenter and kept in place for a certain time t0 causing an imprint
with a depth of e0. A load F1 is then added to F0 for a time t1 and the imprint reaches a depth of e1.
Load F1 is then removed, leaving load F0. The imprint then has a depth e2, which is less than e1 but
greater than e0 due to plastic deformation.
The difference (e2 – e0) indicates the hardness. This difference is lower in harder materials.
The most important scales are the B and C scales, obtained by using spherical and conical
indenters, respectively.
Fig. 26
Microhardness – Portable Durometers
18
Microhardness tests can be used to determine the hardness even of individual grains, and are
carried out on smooth polished surfaces. Either a metallography microscope with very small
indenters or special ultrasound instruments are used (Fig. 27 & 28).
Portable durometers can be of the spring-loaded type (although these are seldom used as the
spring loses its calibration) or of a POLDI type.
With this latter type the load is applied by striking a punch with a spherical end with a hammer.
Between the sphere and the part that is struck there is a metal bar of known hardness. The sphere
therefore causes an indent in both the bar and the test piece and these are used to determine the
hardness of the test piece, with the help of hardness conversion tables, without the intensity of the
blow playing any part, as this may be arbitrary.
UCI_E.exe
dynapocket_rebound.exe
Fig. 27 Fig. 28
Toughness
Toughness can be defined as a metal’s capacity to prevent breaking when it is subjected to
dynamic loads and therefore in an unfavourable state for absorbing the energy of plastic
deformation.
In fact, experience has shown that a material considered as being ductile after normal tensile or
bending tests may act in a fragile manner (that is, without toughness, breaking with very little
deformation) when it is subjected to dynamic forces other than those used for the two tests referred
to above.
It has been found that under certain
conditions a ductile material may not be
tough, while generally material that is not very
ductile is not very tough either.
The toughness of a metal is influenced by
three factors:
- the speed with which the load is
applied;
- the type of force applied (mono-
axial or multi-axial);
- the temperature of the metal (Fig.
29).
Fig. 29
If a ductile metal is loaded quickly, applying a multi-axial force and at low temperature, it may
react with limited toughness.
Where possible, these conditions are simulated in a resilience test, which, in terms of simplicity and
ease of execution is the test most commonly used for determining the toughness of materials.
This test is based on the impact of a weight against a test piece with a notch acting as a stress
raiser (the notch causes a localised increase in the stress and transforms mono-axial stress into multi-
axial stress). This is done at a certain temperature that relates to the working temperature for the
material being tested.
During the test the weight made up of a pendulum hammer strikes a test piece (55 x 10 x 10 mm)
resting on two supports 40 mm apart. The energy (in Joules) absorbed by the test piece to break it
is measured (Fig. 30 & 31).
19
The test piece most commonly used for resilience tests is known as a Charpy V test piece, where
the notch is V-shaped with an opening angle of 45°, depth of 2 mm, and radius at the bottom of
the notch of 0,25 mm.
The size of the grain is very important for good resilience. The smaller the grain, the greater the
toughness.
Generally, it is not possible to predict how a material will behave at a temperature even slightly
lower than the test temperature, due to the tenacity transition phenomenon that we will look at
when dealing with brittle fracture in welded structures (Fig. 31a/b).
Fig. 30
Fig. 31
Errore. Il collegamento non è valido.
20
Fig. 31a - Carbon steel
Fig. 31b – austenitic stainless steel
21
3. Steel Metallurgy
Ferrous-Carbon State Diagram
The Fe-C diagram is a graphic representation of the allotropic transformations that take place in
relation to temperature (Fig. 32).
Fig. 32
It can therefore be used to identify the range over which the various steel structures that can be
found in steel exist, bearing in mind the C content and the steel temperature.
The indications that can be taken from this diagram only apply for very slow heating and cooling.
The percentage C content is shown on the abscissa (x-axis) and the temperature on the ordinate
(y-axis).
The diagram clearly shows that when the C content is below 2% we are dealing with steels, while
above this percentage we are dealing with cast irons.
Steels with a C content below 0,8% are said to be hypoeutectoids and those between 0,8 and 2%
are referred to as hypereutectoids.
While pure Fe solidifies at a temperature of 1539°C, Fe-C alloys solidify over a range that goes
upwards as the C content increases. The “liquid” and “solid” lines indicate the start and finish
temperatures for solidification for each alloy.
We also find that, while allotropic transformation of pure Fe takes place at a constant temperature
of 910°C, in the Fe-C alloy this transformation takes place gradually over the interval between lines
A3 and A1.
A3 indicates the boundary between the austenitic and austenitic-ferritic fields and shows the
temperature at which austenite and ferrite are balanced in the case of hypoeutectic steels, above
which only austenite is stable and below which ferrite is found.
A1 indicates the temperature at which pearlite forms during cooling for steel with a C content
greater than or equal to 0,025% and this is constant where the C content varies.
22
To explain the meaning of the various lines and to comment on the phenomena they involve, let’s
look at what happens when an Fe-C alloy with a C content of < 0,80% (hypoeutectoid) cools,
starting from a temperature above its melting temperature (Fig. 33).
Fig. 33
Starting from the top we come across the liquid line where solidification starts with the formation of
the first austenite cells. Over the interval between the liquid curve and the solid curve, solidification
continues with the gradual disappearance of the liquid. When curve I-E is reached, solidification is
complete and nothing else occurs until we come to line A3 (curve G-S).
At this point the transformation from austenite to ferrite in the solid state begins (gamma – alpha
transformation).
Since ferrite dissolves only a limited amount of C, the residual austenite gradually becomes richer
and richer in C, following line A3 until it reaches line A1 where ferrite grains (with C = 0,025%) and
austenite grains (with C = 0,80%) are found together.
On line A1 at a temperature of 723°C (length P-S), the residual austenite is transformed into pearlite,
which is a eutectoid compound made up of very thin alternate parallel laminations of ferrite and
cementite.
Upon further cooling and until ambient temperature is reached, nothing further occurs if we
overlook the very small amounts of tertiary cementite deposited as a result of the decreasing
solubility of C in alpha Fe.
At ambient temperature the micrographic structure is made up of grains of primary ferrite, derived
from the austenite over the A3 - A1 interval, and grains of pearlite that formed at the eutectic
temperature, each made up of thin laminations of juxtaposed ferrite and cementite.
23
Fig. 34 shows the different structures of a
eutectoid steel: starting from the top, the
alloy follows the same structural
transformations as the previous case until it
reaches the eutectoid at 723°C (point S).
Here, the austenite will become unstable and
will be transformed completely into grains of
pearlite formed by alternate parallel
laminations of ferrite and cementite. This
structure remains stable until ambient
temperature.
Fig. 34
Fig. 35 shows the different structures of a
hypereutectoid steel: in this third case as
well, the alloy follows the same structural
transformations as the previous two until it
meets curve ACM (curve S-E).
Under this curve, the austenite becomes over-
saturated with carbon, which leaks out of the
lattice in the form of cementite. As the
temperature drops, the percentage of
cementite increases until it meets the straight
line at 723°C. At this point the residual
austenite (with C=0,80) is trasformed
completely into pearlite.
After cooling, the steel will be characterised
by a mixed cementite-pearlite structure.
Fig. 35
Structural Transformations in Steel A study of the transformations in steel in a state of equilibrium represented by the ferrous-carbon
diagram does not provide any information on the actual transformations that take place upon
cooling, which are closely linked to the rate of cooling and thus the time over which this takes
place.
During cooling below A3, austenite (Fig. 36) decomposes, giving rise to two processes:
- transformation of the centred face lattice of austenite into the centred body cubic lattice of
ferrite (Fig. 37)
- separation of the carbon from the austenite, in the form of cementite (Fig. 38).
This transformation of the lattice takes place easily as the two lattices relate to one another in
simple geometric terms.
24
Fig. 36 Fig. 37 Fig. 38
As far as the separation of carbon is concerned, as the rate of cooling increases, an unstable
equilibrium is produced in the material with a phenomenon of under-cooling. As this increases, the
forces that tend to bring about a change in the lattice grow. The austenite lattice becomes more
and more unstable and the carbon it contains is in an ever-increasing state of over-saturation.
The delay in the transformation due to the fast cooling speed resulting in over-saturation of
austenite leads to the formation of structures that are more and more fine.
If this speed increases, bainite appears. It has an acicular structure with carbides that are finely
dispersed in the mass.
“Superior” bainite (Fig. 39) is formed due to transformation at higher temperatures. It is made up of
ferrite and carbides in a rounded shape and of “slabs” that are rather coarse, with low hardness
and toughness.
At higher cooling speeds the structure is modified, the temperature at which it is formed drops, and
the ferrite and carbides take on an acicular shape. The “inferior” bainite (Fig. 40) that results has a
very fine structure with good toughness characteristics.
Fig. 39
Fig. 40
Fig. 41
25
If the rate of cooling is increased further, the
carbon that does not find a place for itself in
the centred body lattice that forms is unable
to organise itself in order to bring about a
second phase. It therefore remains in an
over-saturated solution causing a distortion of
the transformed cubic lattice, which
becomes tetragonal (Fig. 42).
This metastable phase is known as martensite
(Fig. 41). It is extremely hard for two reasons:
the tetragonal lattice does not have all the
combinations of sliding planes (and therefore
the relative ductility) that the cubic lattice
has, and then the carbon that remains
trapped in the interstices impedes sliding. The
hardness and fragility of the martensite
increase as the carbon content increases.
Fig. 42
TTT and CCT Curves The temperatures at which the transformations dealt with in the previous point occur and the
speed of cooling, which is critical for these transformations, depend on the type of steel and
especially its chemical composition.
Numerous studies have been conducted into the structural transformation processes that take
place during cooling and the TTT or CCT curves that have been drawn for the various types of steel
can be used to predict the final structures in steel in relation to the heat cycle it has gone through.
The TTT (temperature-time-transformation) curves can be used to predict the structures that will
result from isothermal transformation, i.e. the transformation that takes place when steel is cooled
quickly from a temperature above that at which austenite forms to a lower temperature, remaining
at this temperature for a certain period of time.
The CCT (continuous cooling transformation) curves indicate the structures that result from
anisothermal transformation, i.e. the transformation that takes place when steel is cooled
continuously, starting from a temperature above that at which austenite is formed. These curves
are more commonly used, as they are more representative of real situations such as a welding heat
cycle.
A CCT diagram for continuous cooling of a hypoeutectoid steel is shown in a simplified form in Fig.
43.
The green lines indicated on the diagram plot some curves that relate to a certain number of
cooling speeds.
In this diagram you can see that for slow cooling rates, and therefore bland heat cycles, the lines
marking the start of transformation into ferrite and pearlite and the end of transformation are
crossed. In this case the ferritic-pearlitic transformation has the time to be completed and the final
structure will be of a normalised grain type.
For more severe heat cycles, i.e. for cooling speeds with hardness value of 20 HRC, the end of
transformation line is no longer crossed, and bainite and a certain amount of cementite start to
form.
For cooling speeds that are even higher, such as those shown by curves with hardness value of 50
HRC, the final structure is totally martensitic and the steel attains its maximum hardness.
In this type of diagram the percentage of the structure that will form, which can be detected by a
final macrographic examination, is indicated at the point at which each curve leaves the
transformation field.
The hardness of the final structure is normally indicated inside a circle at the end of the curve itself.
26
Fig. 43
27
4. Manufacturing and Classification of Steels Methods of Steel Production, Refining, Casting and Lamination
Production of Pig Iron:
Pig iron is made by smelting iron ore with coke in a blast furnace.
Together with the coke and iron ore, limestone is put in the furnace (Fig. 44 & 45).
The blast furnace is used to produce pig iron (iron-carbon alloy) by reducing the metal oxides (for
example Fe2O3) naturally present due to the oxidising atmosphere. The oxygen bonds with the
carbon added to the material to be produced, forming carbon dioxide and leaving the pig iron.
The limestone purifies the pig iron in the bath, bonding with the impurities present and forming a
slag that floats on the surface and out of the iron.
The pig iron that is extracted from the blast furnace can be further refined for steel-making.
The heat for production of pig iron is generated by combustion of the coke mixed with the iron ore.
The combustion air is preheated inside special towers which are built into the plant. The blast
furnace is used for continuous production throughout its service life (several years), which normally
depends on the durability of its inner lining.
Fig. 44 Fig. 45
28
Steel Production: It is possible to distinguish three manufacturing phases in steel production:
- the first phase covers the actual production of molten steel (refining and processing)
- the second phase covers the casting of the steel from the ladle in ingots or continuous
castings
- the third phase covers the lamination of the semi-finished goods (ingots, slabs, billets).
Refining: In brief, refining is the oxidisation process used to eliminate excess components (carbon and silicon
in the cast iron) and dangerous impurities from the initial bath to produce a product that is as pure
as possible and that contains only the required quantity of carbon, manganese and silicon.
This process involves an initial load of molten, solid or mixed items in converters (Fig. 46 & 47) or
electric furnaces (Fig. 48).
Converters use a load of molten cast iron produced in blast furnaces as their main constituent,
whereas electric furnaces use a solid load of scrap recycled steel with cast iron and/or graphite
added.
The chemical reactions that occur during the refining process can be summarised as:
• Decarburisation: the injection of oxygen causes an initial oxidation reaction in the iron and
silicon. The oxidation reaction in silicon is strongly exothermic, raises the temperature of the
bath and initiates the conversion of the carbon to carbon monoxide (CO) and carbon dioxide
(CO2), which are released as gases.
• Deoxidation: during the decarburisation reactions, manganese is also oxidised and reacts with
the silicon and iron oxides to form silicates that pass into the slag and deoxidise the molten
bath.
• Purification: the manganese and calcium and magnesium oxides or carbonates introduced
into the bath react with sulphur and phosphorus compounds to form manganese sulphides and
calcium phosphates that pass into the slag with consequent purification of the bath.
The steel produced in this way in the converter or electric furnace still contains notable quantities of
dissolved gases, particularly oxygen, which must be eliminated for deoxidation to be completed in
the ladle with the addition of elements such as manganese, silicon and aluminium that react with
the oxygen to form oxides that precipitate in the metallic matrix.
Steels are divided into three categories on the basis of the degree of deoxidation, i.e. the level of
oxygen remaining in solution in the molten product:
• Effervescent steels: these have a high oxygen content that causes notable development of
gaseous carbon monoxide during casting and solidification due to the FeO + C = Fe + CO
reaction which, in turn, causes the development of a continuous layer of blowholes in the
peripheral zones, particularly the upper peripheral zones. The outer surface layer of the ingot is
made up of purer, more compact metal with a low carbon content.
• Semi-killed steels: these do not undergo deoxidation to the extent of blocking the reaction
between the carbon and oxygen. They therefore have a structure with less porosity and fewer
blowholes than effervescent steels and a smaller central segregation.
• Killed steels: these undergo forced deoxidation by the injection of deoxidising elements such as
silicon and aluminium to block the development of other decarbonising reactions and
consequent effervescence during solidification. After solidification these steels have a
compact structure without blowholes but with a shrinkage cone at the top of the ingot and
that must be removed before lamination. The addition of aluminium or other elements such as
vanadium, niobium or titanium causes further refining of the grain, producing fine-grained killed
steels.
The LD Converter
The LD process is the most commonly used in steel production and uses molten cast iron and a cold
load of scrap and solid cast iron.
The oxidising agent used in the process is pure oxygen.
The LD process was introduced in the 1950s and takes its name from the Linz and Donawitz
Steelworks in Austria, which first used it.
29
The converter consists of a cylindrical steel container with a convex base, over which a truncated
cone-shaped part is fitted.
The converter is lined with dolomite or magnesite fettling and a basic slag is formed by the
introduction of fluxes such as lime and fluorine.
The converter is loaded in an inclined position and then the oxygen injection nozzle is
concentrically positioned above the opening for the injection of the oxygen (Fig. 46 & 47).
Fig. 46 Fig. 47
During the injection, which lasts about 20 min and is calculated on the basis of the oxygen volume
necessary for the desired final carbon percentage, silicon is the first element to be oxidised,
followed by manganese. These begin to form a slag with a certain basicity and decarburisation
begins. When this phase is concluded, desulphurisation begins and the slag reaches its maximum
basicity.
When the required quantity of oxygen has been injected, the converter is inclined and a sample
taken for analysis. If this shows the required analytical values for the principal chemical elements
(carbon and manganese) and the impurities (sulphur and phosphorus), iron alloys are added for
deoxidation and correction of the casting, which is finished in the ladle.
The Electric Furnace
Electric furnaces are either indirect arc (the bath is heated by radiation from an arc that develops
between electrodes above it) or direct arc (the arc is formed by electrodes and the bath and the
electric circuit is enclosed in the bath).
Depending on the type of steel required, the furnace (Fig. 48 & 49) is loaded with scrap as well as
with carbon-bearing material (coke, graphite powder, cast iron) and lime, feldspar and iron
mineral to form the slag.
The electric current from the electrodes melts the load with the aid of the injected oxygen. The
temperature of the bath and its surroundings is increased until the scrap is completely melted at
temperatures strictly correlated with the carbon content of the molten bath.
The refining of the bath begins with an oxidation phase, during which the injection of oxygen helps
decarburisation and the metal-slag interchange, above all favouring dephosphorisation of the
bath: the phosphorus that has accumulated in the slag is eliminated during the first scorification.
The slag is renewed with the addition of lime, feldspar and reduction materials such as ferrosilicon,
calcium silicide and aluminium. These deoxidisers reduce the iron oxide and the other metal oxides
in the slag, and during this reduction phase desulphurisation of the bath takes place.
30
Fig. 48 Fig. 49
Iron alloys are added to obtain the required chemical balance and the casting is then poured into
the ladle.
By means of the electric furnace it is possible to produce a wider range of special steel products
than by the LD process: from unalloyed steels to high alloy steels. The electric furnace is particularly
irreplaceable in the production of alloyed and special steels as it is possible to obtain the necessary
content of the alloying elements by controlling their oxidation process.
Steel Casting
The casting systems used in steel manufacturing are:
- casting in ingots that can be transformed into slabs and billets by cutting and laminating
- continuous casting in slabs and billets.
Casting in lngots
This is the oldest casting process and now almost completely superseded by continuous casting.
However, it remains the only technique for casting non-deoxidated steels (rimmed and semi-killed
steels).
Furthermore, ingots are still used for the production of forged steels of particularly large dimensions.
The steel is cast in cast iron moulds with walls of great thickness to enable a rapid solidification of
the molten material.
The moulds are generally in the form of truncated pyramids, with square or rectangular sections
and with the largest side uppermost to make it possible to easily release the ingot by pulling it
upwards.
The most commonly used method is direct casting: the steel mould that emerges from the bottom
of the ladle falls freely into the ingot (Fig. 50).
Another casting method used when a large number of ingots of small dimensions is required is
bottom casting: the steel is forced upwards into the mould from below, using a system of interlinked
canals made of refractory brick (Fig. 51).
31
Fig. 50 Fig. 51
The solidification of the cast steel in moulds begins in the surface layers adjacent to the internal
walls of the mould and proceeds in the opposite direction to the heat flow, i.e. perpendicular to
the walls themselves.
When solidification occurs, it is possible to distinguish three different zones in a cross-section of the
ingot:
• the skin zone, which consists of very fine-grained steel due to the very rapid solidification of the
metal in contact with the mould walls
• the intermediate zone of columnar crystallisation, which consists of dendritic crystals elongated
towards the centre and perpendicular to the walls of the mould
• the central zone, which consists of globular crystals with equal axes.
Due to the effects of the different rates of solidification, the steel components with the lowest
solidification point (generally impurities) are concentrated in the central part of the ingot, which
solidifies last, and cause the phenomenon of macro-segregation.
Furthermore, due to the effect of contraction in volume that occurs in the course of cooling, and to
a greater extent with solidification, a shrinkage cone is formed in the centre of the ingot with
consequent problems for its subsequent use.
To reduce the size of the cone and its harmful effects, a supplementary generally insulated part,
called the head, is placed above the mould to thermally isolate the upper part and slow the
cooling process, thereby producing a shallower shrinkage cone and collecting the bulk of the
impurities and oxides in that zone, which can then be removed.
Continuous Casting
The continuous casting system is actually the most commonly used due to its manufacturing and
economic advantages (less material lost and elimination of the primary lamination phase in
converting ingots into slabs and billets).
The possibility of using continuous casting depends on the maximum dimensions of the slabs and
billets that can be produced, and also on the final thickness required for the finished product. In
fact, to obtain a good, homogenous finished product it is always necessary to have a reduction in
the thickness of the rolling mill of at least 6 to 1.
The continuous casting process can be summarised as follows (Fig. 52-53-54).
The molten steel in the ladle is poured into an intermediate container called a tundish, which
controls the steel flow into the machine and, if necessary, divides it between the different tubes
that make up the machine.
The steel is poured from the tundish into bottomless copper moulds with a vertical oscillation
mechanism and is cooled by circulating water. A regulating device, called a false slab, is inserted
into the mould to serve as the base before the casting.
32
The oscillating mould, whose length is calculated to permit the formation of a thin skin layer to
enable the steel to support itself, accompanies the solidifying bar in its descent towards the
bottom.
Below the mould there are casting, extractor rollers that begin to pull the false bottom away just
seconds after casting begins, permitting the semi-finished product to move inside the mould; the
steel begins to solidify at its contact points with the mould walls, assuming the mould form and
finishing the solidification process in an underlying zone where it is subjected to jets of water.
Under the extractor rollers the half-product goes through an oxygen-acetylene station where it is
cut into the required lengths.
Fig. 52
Fig. 53 Fig. 54
There are three types of continuous casting machines:
- vertical machines
- vertical machines with bending
- curved machines
The steel solidifies in the same way as that described for mould casting.
33
Due to the continuous feeding of the molten steel, shrinkage cones cannot form and less material is
wasted.
However, as there is less waste material with its concentration of impurities, it is necessary to pay
particular attention to the phenomenon of macro-segregation, which tends to concentrate all the
impurities in the centre of the bar and could cause delamination of the laminated product.
To avoid this phenomenon, electromagnetic stirring in the mould has been introduced by placing
an electric winding around the mould. In this way a better distribution of the crystallisation nuclei is
obtained with a subsequent reduction in macro-segregation and the creation of a structure with
more-or-less equal axes for most of the thickness.
Furthermore, the steels used have a low level of carbon or impurities.
Rolling
The rolling process aims to elongate the product by a reduction imposed on the transverse section
using pressure rollers that form a lamination train (Fig. 55 & 56).
The products obtained by rolling are the following:
- plates: hot laminates with thickness ≥ 3 mm
- sheets: hot or cold laminates with thickness < 3 mm
- long products: section bars, rails, rods, etc
- pipes
In this article we will only deal with plates.
A plant for the production of plates consists of:
1. heating furnaces
2. rolling mill
3. hot levellers
4. hot cropping shears to trim the ends of the plates
5. cooling beds and control
6. edging shears to trim the sides of the sheet
7. shears to cut the product to the required dimensions.
Fig. 55 Fig. 56
Lamination trains can differ according to the number and dimensions of the lamination cylinders:
two, three, four.
34
The most commonly used rolling mill for the
production of plates is the reversible four-roller
one (Fig. 57).
This consists of four cylinders vertically
positioned one on top of the other. Only the
two middle, work cylinders are operated by
the motor and are smaller than the other two
on which they rest. This arrangement is based
on the principle that in lamination, the rate of
elongation increases as the diameter of the
cylinders decreases, which makes it possible
to impose high pressure because the stress
imposed on the working cylinders is absorbed
by the robust support cylinders. It is called
reversible because it can work in two
directions.
Fig. 57
Rolling Methods
There are three types of rolling processes (Fig. 58).
• Standard Rolling: used when high mechanical characteristics, especially resistance, are not
required and it is desired to contain the costs of production. This process usually finishes at quite
a high temperature to reduce the energy required in the hot plastic lamination deformation. As
the process cannot be controlled there could be a certain variability in the characteristics of the
product.
• Normalising Rolling (also called controlled rolling): the rolling process in which the final
deformation occurs within the temperature interval for standard thermal normalisation
treatment so that the structure of the material is generally the same as that obtained with
normalisation. The required values for the mechanical characteristics are kept even after any
subsequent normalisation treatment.
• Thermo-mechanical Rolling: the lamination process that provides strict control of the
temperature of the plate and the degree of lamination. In general, most of the lamination is
obtained around the temperature of Ar3, which means to say that it could cause the finishing of
the lamination towards the lower part of the temperature interval of the inter-critical bi-phase
zone. The properties conferred by thermo-mechanical rolling are not maintained in the event of
a successive normalisation treatment or another thermal treatment.
35
Fig. 58
Recrystallisation
The different structures and characteristics coming from the different rolling processes are
determined, for the most part, by the recrystallisation processes.
During the high-temperature rolling, austenite grains are strongly deformed, fractured and
elongated by the lamination rollers. As already stated, due to the high temperature and energy of
the stored deformation, the material tends to recover the state it had previously and to
“recrystallise”.
At a higher temperature and/or greater deformation, recrystallisation occurs during the
deformation phase itself (dynamic recrystallisation): due to the very high temperature the
austenitic grains grow back after recrystallisation, although they do not regain their original
dimensions.
At a lower temperature or a weaker deformation, recrystallisation only occurs at the end of the
deformation phase (static recrystallisation) and the dimensions of the recrystallised grains are
smaller.
If rolling occurs at an even lower temperature, recrystallisation does not occur, or only partially, and
the austenitic grains remain deformed.
The temperature for dynamic or static recrystallisation, partial recrystallisation or non-
recrystallisation basically depends on the degree of deformation and the type of steel. Certain
micro-alloying materials such as niobium slow recrystallisation: that is, they increase the
temperature or the temperature time interval and so recrystallisation may occur.
The dimension of the ferritic grain obtained after the transformation is conditioned by
recrystallisation. In fact, the nucleation of ferrite begins at the edges of the recrystallised austenitic
grains and proceeds towards the inside of the latter. Instead, in deformed austenitic grains the
36
nucleation of the ferrite not only takes place along the edge of the grain but also along the
deformation band formed inside the grains, producing a finer ferrite grain.
Continuing the rolling after the beginning of the austenitic transformation (gamma-alpha field)
further deforms the residual austenitic grains and deforms and fractures the ferritic grains already
formed, forming sub-grains and, consequently, a very fine final structure. Standard Rolling
As already said, standard rolling is carried out at a sufficiently high temperature to cause complete
recrystallisation of the grains and their subsequent growth. As it is not controllable, it can cause a
certain variability in the characteristics of the product in the same laminate and, more probably,
between rolled products of the same casting and thickness.
This process is used for steel plates of ordinary quality such as hull structural steel - A and steels for
high-temperature uses.
Normalising Rolling
Normalising rolling permits control of the temperature of the last rolling pass and the thickness
reduction and lamination cycle methods immediately preceding the last pass.
The final rolling pass is performed in a completely austenitic field but just above the gamma-alpha
A3 transformation line so that only static recrystallisation of the structure occurs, without an increase
in the grain or only a partial recrystallisation of the deformed grains. In this case it is important that
the final rolling passes are sufficiently strong (high percentage of thickness reduction) to affect the
entire thickness and avoid the formation of a mixed structure of large grains of undeformed
austenite and small grains of deformed and partially recrystallised austenite.
Thermo-mechanical Rolling
The rolling is carried out in two cycles:
- a first (roughcast) cycle at high recrystallisation temperature and, after a suitable waiting
period,
- a second (finishing) cycle at a lower “non-recrystallisation” temperature that can also
conclude in the gamma-alpha interval.
This method is based on the following principles:
• to benefit from the formation and presence of carbonitriding materials, in particular Nb, but
also other micro-alloying materials (V, Ti, B) that are so stable at high temperatures that they
can block the growth of austenitic grains during the recrystallisation phase during the waiting
period following the first cycle and during the second lamination cycle.
• to perform a part of the rolling at a temperature at which austenite does not recrystallise
further and make the ferrite in the austenite grain nuclear at the “deformation” band with the
consequent formation of a finer ferritic grain. The presence of Nb in the form of carbonitriding
materials in the non-precipitated part at the edge of the grain and still dissolved in the
austenite raises the “non-recrystallisation” temperature.
• lastly, to carry out a part of the lamination in the austenitic-ferritic zone in such a way as to
crush the ferritic grains into “sub-grains”.
Due to the high-temperature refining of the grain, which is not possible with the normalising process,
this type of rolling makes it possible to obtain products with high resistance and toughness.
However, as these characteristics are the consequence of thermo-mechanical treatment, which is
the result of a delicate equilibrium between the crystalline structure obtained (extra-fine grain) and
the carbonitriding materials precipitated and their state of dispersal in the structure, the product is
very sensitive to actions that can alter or destroy that equilibrium, especially prolonged high
temperatures for thermal treatments, hot moulding and high-temperature welding.
In general, the characteristics of plates produced with thermo-mechanical procedures are
guaranteed only for heat treatments at a temperature not exceeding 580°C.
37
Classification of Steels
Steels can be classified as follows:
- carbon and carbon-manganese steels;
- low alloy steels, i.e. steels containing up to 5% of alloying elements;
- high alloy steels, i.e. steels containing over 5% of alloying elements.
Normally, high alloy steels are stainless steels or special steels, produced for specific purposes,
whereas carbon steels and low alloy steels are employed as construction steels and for general
uses.
Carbon and carbon-manganese steels
These are the most commonly used and the least highly regarded; they are easily worked and
readily weldable. However, they are subject to corrosion, above all at high temperatures, at which
they also lose mechanical resistance, while at very low temperatures they become brittle.
They are divided into mild steels, with carbon content less than 1%, and hard steels, with higher
carbon content: mild steels are very ductile and malleable, they are easily worked and have
excellent tensile and compressive strength; they are also very resilient.
Ordinary and high strength steels for hulls belong to this family of steels.
Hard steels are less resilient, that is they are more likely to fracture if subjected to violent impact, but
they also have much greater surface hardness. Also, they are very good for hardening, unlike mild
steels. On the other hand, they are obviously less easily worked.
Quenched and tempered high strength steels
These are normally low alloy steels whose chemical composition is richer than ordinary C-Mn steels
due to the addition of elements (Cr, Ni, Mo, V, Nb, B) favouring their hardening. The quenching
and tempering to which they are subjected substantially increases their tensile strength and
toughness.
Chromium-molybdenum steels
These steels are suitable for the construction of boilers and pressure vessels operated at elevated
temperatures. The chromium and molybdenum, which are present in the solid solution of the ferrite
matrix or in the form of carbides, reduce the tendency of the structure of the steel to deteriorate at
high temperatures and therefore increase creep resistance. Also, the chromium provides improved
oxidation resistance at elevated temperatures.
Ferritic steels for low temperatures
The following techniques are adopted to improve the performance at low temperatures of C-Mn
steels:
- complete deoxidation and denitrification of the steel and uniform distribution of the impurities
so as to avoid local concentrations of segregations and improve the ductility and toughness of
the steel
- attainment of structures with a fine grain by means of heat treatments or thermo-mechanical
treatments.
For low temperature service less than –60°C, nickel additions of increasing amounts (from 1 to 9%)
are used as a function of the minimum operating temperature. The nickel shifts the ductile to brittle
transition curve towards low temperatures without causing an excessive increase in hardness.
These steels are characterised by bainitic/martensitic structures with levels of residual austenite up
to 20% in the case of steel containing 9% Ni.
38
Stainless Steels
Stainless steels are Fe-Cr or Fe-Cr-Ni type ferrous alloys with chromium content from 12% to 30% and
up to 25% nickel. Chromium, which is a less noble metal than iron in terms of electrochemical
potential, oxidises forming a tough continuous layer which is impervious to water and highly
resistant to many corrosive agents, thereby protecting the metal beneath. The film self-repairs in the
presence of oxygen if the steel is damaged. In the context of corrosion prevention, the
spontaneous formation of this hard, non-reactive surface film is known as passivaton.
Stainless steels can be divided into the following families:
� Austenitic and superaustenitic steels containing Cr-Ni or Cr-Mn-Ni
� Martensitic steels containing Cr (11-16%) and Ni (max 4%)
� Ferritic steels containing Cr (11-25%)
� Austenitic-ferritic (duplex) steels containing Cr (18-28%), Ni (4-9%) and Mo (1,5-3%)
� Precipitation-hardening steels.
39
5. Heat Treatment for Steel
Heat treatments vary in terms of the temperatures reached and the way they are carried out, and
are generally broken down into two categories:
- treatments at temperatures above A3: annealing
normalising
quenching
- treatments at temperatures below A1: tempering
stress relieving
Annealing
This heat treatment is carried out by heating the material to above A3. The initial micrographic
structure is transformed into an austenitic structure; the less the temperature is pushed beyond A3,
the finer this structure remains. Steps must be taken, however, to make sure that the austenitic
transformation takes place throughout the entire metal mass.
Slow cooling in the furnace can then take place so as to cross the A3 and A1 lines over a period of
sufficient length to obtain complete separation of ferrite and pearlite (Fig. 59).
A heat treatment of this kind can be used to obtain a final micrographic structure that is of a
granular type with equal axes. In this way the coarse solidification structures are eliminated, these
being more fragile, or less plastic, than the transformation structures.
In addition, if the time spent in the austenitic range is sufficiently long, annealing can be used to
retain the differences in composition, constitution and structure using diffusion processes, and to
eliminate structural distortions due to plastic working for example.
Fig. 59
Normalising
Normalising is a heat treatment that differs from annealing in terms of the greater cooling speed
used.
After the time spent in the austenitic field the piece is cooled in still air rather than in the furnace.
The A3 – A1 interval is passed through more quickly, giving rise to a ferrite + pearlite structure that is
still granular but finer (Fig. 60-61a/b-62a/b).
Grain fineness, which is typical of normalised structures, generally gives the metal better
mechanical properties, as we have already seen.
40
Fig. 60
Normalising is also intended to transform a coarse and uneven structure into a homogeneous
structure with fine grains: the figures below show examples of the effects of normalising a casting
and a plate.
If the C content is rather high (over 0,30), however, cooling in the air may give rise to initial
hardening phenomena with a significant reduction in ductility. In this case annealing is preferable.
Fig. 61a – coarse casting Fig. 61b normalised casting
Fig. 62a – plate in the “as rolled” condition Fig. 62b normalised plate
Quenching
Quenching is a heat treatment that is carried out by heating the material to a temperature
exceeding A3 and then cooling it very quickly in water or oil (Fig. 63).
41
The quick cooling does not allow time for austenite to be transformed into ferrite and pearlite and,
as already indicated in previous paragraphs, bainite type structures are formed or, in the case of
very fast cooling, hard, fragile martensitic structures are formed.
Fig. 63
There are many factors that influence the hardening characteristics of a material, including:
o cooling speed
o percentage of carbon
o presence of alloying elements
o grain size
o temperature at which austenitic transformation takes place
o presence of unmelted elements
o heating speed.
As we have already said, the cooling speed is a basic factor and as this increases so does the
probability of producing martensitic structures.
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Carbon and alloying elements move the TTT or CCT curves to the right, and so for the same cooling
rate and other factors, they favour the formation of hardened structures.
Grain size plays an important role. In fact, ferritic-pearlitic transformation comes about due to
nucleation at the edge of the grain and, bearing in mind that everything that inhibits this
transformation favours hardening, the larger the grain is the less space there is available around the
grain, and the more easily hardening will occur.
The higher the temperature at which austenitic transformation occurs, the more the grain is
overheated and the larger its size, which makes the steel easier to harden.
Elements that have not melted gather around the edges of the grains and set up a centre for
germination for pearlitic transformation, thereby limiting the steel’s susceptibility to hardening.
Very quick heating and a very short time in the austenitic-ferritic field (between A1 and A3), as
occurs in welding, may lead to a low diffusion and homogenisation of the carbon released by the
cementite, thereby creating zones that are rich in carbon. When these areas are then cooled,
they may tend to harden even when the average composition of the steel would exclude this
possibility.
42
Apart from the rate of cooling, C content and that of other alloying elements are also very
important factors.
There is a formula that includes the various elements and their influence on the hardening capacity
of a material. This is the C equivalent formula:
Ceq= C + Mn/6 + (Cr+Mo+V)/5 + (Ni+Cu)/15
Hardening brings about an increase in the hardness of the material and significant reduction in
ductility and toughness. It is therefore very important to try to avoid hardening phenomena,
especially following welding operations.
In some cases hardening treatment is applied voluntarily. This is the case in hardening and
tempering that is used to obtain materials that have notable mechanical characteristics,
associated with good toughness characteristics.
Quenching and tempering involves the combination of two heat treatments: hardening followed
by tempering (Fig. 64).
Fig. 64
Tempering
Let’s look at the action of tempering treatment, which is carried out at temperatures of between
400 and 650°C.
At these temperatures the acicular structure of the martensite (Fig. 65 A) is transformed into a non-
laminar structure of ferrite and cementite (Fig. 65 B).
Fig. 65
Hardness is greatly reduced and, what is more, the higher the tempering temperature the more the
steel’s ductility and toughness increase (Fig. 66 & 67).
43
Fig. 66 Fig. 67
This is caused by the carbon escaping from the distorted tetragonal structure of the martensite,
while the grain size remains unchanged and is very small due to extremely quick cooling.
Stress Relieving
Stress relieving treatment is essentially intended to eliminate or rather reduce as much as possible
the residual tension that is an inevitable result of any welding operation.
For steels commonly used in construction this treatment is carried out at 600 – 650°C, with this
temperature generally being maintained for two minutes for each millimetre of thickness, and a
minimum of 30 minutes (Fig. 68).
The efficacy of the treatment is due to the fact that at high temperature the yield strength of the
material is reduced to values that are practically negligible.
Thus, when the entire welded structure is heated to this temperature and kept there sufficiently
long, the tension is released and reduced to the value of the yield strength at that temperature.
This release of tension may result in plastic deformation and so, after heat treatment, the
dimensions of the structure are changed to a greater or lesser extent.
Naturally, since this heat treatment is carried out at the same temperature as tempering, the
beneficial effects of the two are added to one another and thus things such as hardness points
can be eliminated in specific cases.
Fig. 68
44
WELDABILITY
1. Metallurgy of Welds in Steels
An autogenous weld, i.e. a weld involving the fusion of the basic material, produces a joint that has
two distinct zones (Fig. 69):
- the WELD ZONE (W.Z.), made up of the basic metal that has been melted and any filler metal,
i.e. the solidified weld pool
- the HEAT AFFECTED ZONE (H.A.Z.), made up of the basic material brought to a high
temperature in the heat cycle in welding and subsequently cooled.
Fig. 69
The Weld Zone
The structure of the weld zone is greatly influenced by the parameters and the way the joint is
formed, and especially by the number of passes used.
Let’s have a look at these variables in relation to the number of passes.
Welding in a single pass
Despite being quick, solidification of the fusion puddle is not instantaneous, and this influences the
structure of the weld zone, which is strictly dependent on the way it solidifies.
The quickest way of disposing of the heat is constituted mainly by the surface in contact with the
weld zone and the surrounding basic material. This surface forms the basis for the germination of
the grains that develop in the liquid zone.
The orientation of the dendrites is aligned with the direction of cooling, and the austenitic grains
develop by elongating towards the centre of the melted area, forming a columnar dendritic type
of structure (Fig. 70 & 71).
Fig. 70 Fig. 71
The grains start off from those in the HAZ, linking up to these and keeping the same orientation as
the lattice (epitaxy) and the same size. In fact, the larger the grains in the HAZ are, the larger the
dendrites that form in the WZ tend to be.
Later on we will see how grain size and shape affect the likelihood of cracking when hot.
HEAT-AFFECTED ZONE BASE METAL WELD METAL
45
Dendritic solidification of the weld zone is always accompanied by conspicuous phenomena of
segregation, with impurities building up at the edges of the grains, in the areas that solidify last, that
is, at the centre of the joint.
The speed of solidification determines the size and thus the number of dendrites. Slow solidification
produces large grains, while quick solidification gives rise to more slender and numerous dendrites.
The speed of solidification basically depends on the following factors:
- thickness of the pieces: the thicker they are, the quicker they cool.
- welding procedure – the higher the heat build-up, the slower solidification takes place.
From this point of view the most common procedures have the following characteristics:
- Submerged arc automatic welding is characterised by slow solidification and very coarse
dendritic structures. Great power is involved and the heat build-up is rather high despite the
speed with which the weld is formed.
- Although oxygen-acetylene welding uses a heat source that is a lot less powerful than the arc,
it heats the piece up a lot because it moves forward slowly. This results in relatively slow
solidification with rather coarse dendritic structures being formed.
- The heat build-up in arc welding by hand varies greatly in terms of the diameter of the
electrode, the intensity of the current, and the speed at which the weld moves forward. The
bigger the pass, the slower it cools. In general, the dendritic structure is finer than in the
previous two cases.
- The structure left by TIG welding is finer still, involving a low heat build-up and a cooling effect of
the inert gas that is blown onto the welding puddle.
Irrespective of the procedure used, after solidifying, the material in the weld zone undergoes
transformations in the solid phase that depend on its carbon content and that of other alloying
elements, as well as on the cooling speed.
In general, for steels with low and medium carbon content and low alloy content, ferrite-pearlite
structures are obtained of varying grain coarseness and with a composition that is more or less
homogenous, depending on the extent of segregation.
Hardening phenomena can only occur where there is a high carbon content and, more
commonly, with high cooling speeds and coarse austenitic grains.
Welding in a number of passes
When welding is done in a number of passes the heat of the successive passes has a beneficial
heat treatment action on the passes already completed.
In fact, each pass is partially melted and incorporated into the subsequent pass and, at a greater
depth, it undergoes a heat cycle that is similar to that used for normalisation treatment that does
away with the original dendritic structure to replace it with another that is finer.
The phases described and the structures indicated are clearly visible in a buffed etched section
(macrographic examination) (Fig. 72).
The characteristic solidification structure with
its course elongated grains is only found in the
final pass. The other passes are more or less
normalised depending on the heat levels
used and the volume of each pass.
After this process that makes the grains finer,
the mechanical characteristics of the joint
are notably improved and it is more ductile
and tougher.
Fig. 72
46
Multiple pass welds therefore offer the following advantages:
- finer dendritic structures are obtained due to the lower heat build-up for each pass;
- subsequent passes cause a normalising effect;
- phenomena of dilution, segregation and heat cracks are reduced.
Composition of the welded area
In very general terms, autogenous welds are obtained by melting a certain part of the basic metal
and completing the joint by adding a certain amount of filler metal material from outside. The two
molten metals mix in the weld puddle and form a new alloy.
The ratio between the volume of the basic metal and the total volume of the WZ in a welded joint is
known as the “dilution ratio”.
Accordingly, the dilution ratio d (%) is defined as the ratio between the volume of the molten basic
metal and the total volume of the weld zone, multiplied by 100.
Thus, if we call the contribution of the filler metal "Va" and that of the basic metal "Vb", in the case of
a single pass we would have: Rd(%) = Vb/(Va+Vb) x 100 (Fig. 73).
Fig. 73
Dilution is especially favoured by the intensity of the current used and so changes significantly
depending on the welding procedure used. At best, the following values can be obtained:
- Rd%= 0 in brazing
- Rd%= 20 in the first pass for manual arc welding
- Rd%= 10 in subsequent manual arc welding passes
- Rd%= 65 in high penetration submerged arc welding
- Rd%= 100 in resistance welding.
The part played by the basic material in the composition of the weld zone can have a significant
influence on the characteristics of the joint.
In fact, in high-penetration welds the alloying elements and impurities that the basic metal may
introduce into the weld puddle take on particular importance, as they may form alloys that are not
helpful in creating healthy joints.
The influence of alloying elements from the basic material
Due to the limited amounts of special alloying elements in the composition of carbon steels, the
chemical composition and metallurgical and mechanical characteristics of the WZ are mainly
influenced by the percentage of carbon, silicon and impurities (sulphur and phosphorous) in the
weld zone.
Carbon
As we have seen before, a high carbon content favours hardening and so should be avoided as
much as possible in welds.
100×+
=ba
bd VV
VR
Vb
Va+Vb
47
One should also bear in mind that in the WZ this phenomenon is also favoured by the fact that the
austenitic grains are coarse and provide the locations for segregation phenomena that tend to
concentrate the carbon in their marginal zones.
To reduce the risk of martensitic structures, where possible welding is done using low carbon filler
materials to keep the carbon content in the WZ low and below that of the basic metal.
Carbon also tends to make the material less ductile at high temperatures and so, as we will see
later, it favours the occurrence of heat cracks in the WZ.
Finally, carbon can also react with gasses, especially oxygen, or with ferrous oxide, thereby
contributing to the development of porosity where no deoxidising elements are present.
Silicon
Silicon is added to steel to act as a deoxidising agent and is therefore mainly found in killed steels.
It is best to keep the content rather low (not above 0,30%) however, since, as with carbon, it may
have a significant effect in making the WZ fragile.
Sulphur
Sulphur is an impurity that is always found in steels and melting part of the basic metal introduces it
into the weld puddle.
It forms a eutectic with the iron (FeS) that melts at 985°C.
Sulphur has a strong tendency to segregate when the WZ solidifies, accumulating at the edges of
the dendritic grains, where the FeS eutectic remains liquid longer while the surrounding metal is
already solid. This reduces the sections withstanding the shrinkage forces between grains and
favours the phenomenon of cracking when hot.
The sulphur content of the weld zone must therefore be kept to a minimum, and to this end steels
are produced with a very low content that does not exceed 0,04 – 0,05%. In addition, the
presence of sulphur can also cause blowholes due to a reaction with the H2 and O2 absorbed and
dissolved in the molten metal. The action of the sulphur can be attenuated by adding Mn to the
weld puddle, which tends to react with the FeS to form MnS that solidifies in a globular form spread
throughout the metal mass.
In addition, the basic coating on the electrodes have a purifying function and so by welding with
electrodes of this type it is possible to fix the sulphur in the slag in the form of calcium sulphide.
Phosphorous
Like sulphur, phosphorous is also an impurity always found in steel.
It has a strong tendency to segregate in the form of FeP, although to a lesser extent than sulphur, as
phosphorous is soluble in iron to a certain extent, even at ordinary temperatures.
It also further favours cracking when hot, although it does not seem to have a very great effect on
this defect.
Phosphorous segregation also causes fragility at low temperatures, which is therefore mainly felt in
terms of resilience.
The phosphorous content must therefore always be very low (never more than 0,05%). For welding
operations it is generally advisable to use materials that can reduce the phosphorous content in
the puddle (e.g. basic electrodes).
Gas in the weld zone
Some chemical changes and defects in the welded joint can be caused by the molten metal
absorbing gas.
The gasses most susceptible to absorption during welding are oxygen, nitrogen, and hydrogen.
Their action can be seen in phenomena that are at times not yet perfectly understood and that
are generally harmful in welding.
Oxygen
As is well known, oxygen reacts with iron to form oxides that are partly soluble in the iron itself and,
in the case of welding, can result in harmful inclusions.
48
The solubility of oxygen in pure iron at melting temperature is of the order of 0,21% and increases in
molten iron as the temperature increases.
In solid iron this solubility reduces rapidly, so much so that in gamma Fe it is of the order of 0,05%,
and even lower than that in alpha Fe.
Due to this difference in solubility between the liquid and the solid state, oxygen absorbed at high
temperatures tends to leave the puddle as it solidifies, and therefore reacts with the carbon to form
carbon monoxide (CO).
This oxide is the main cause of blowholes, which together with oxide inclusions, impair the
mechanical characteristics of the material.
For these reasons it is important that the puddle be protected against the oxidising action of the air
during welding.
This can be done, for example, by the coating on the electrodes, the volatile nature of which
creates an atmosphere that protects the metal as the arc passes and the weld puddle by means
of the slag that floats on it.
Since a certain amount of oxidation of the iron is inevitable, this action continues due to the
deoxidising elements specifically included in the coating (manganese and silicon) that oxidise or
that reduce the ferrous oxide, thereby avoiding a well oxidised weld puddle.
Nitrogen
As is the case with oxygen, the solubility of nitrogen in iron also decreases with temperature
reductions until it becomes almost non-existent at ambient temperature.
Over-saturated nitrogen therefore tends to form ferrous nitrides that precipitate (normally following
plastic deformation) and cause ageing, resulting in increased fragility and impaired mechanical
characteristics.
In this regard as well, good protection of the weld puddle and of the filler metal during transfer is
very important. This is normally done by the coating on the electrodes or by a protective gas in a
continuous wire procedure for example.
Hydrogen
Due to hydrogen’s particular importance in welding and the complexity of the phenomena
connected with it, see the comments in this regard in the paragraph provided.
We merely wish to state that, as in the previous cases, the solubility of hydrogen is high in molten
steel, low in austenite, and lower still in ferrite, and so during cooling it tends to leave the solution
causing cracks, flakes and porosity in the weld zone and, in the case of steels susceptible to
hardening, the formation of cracks under the bead (cracks when cold).
The Heat Affected Zone
The metallographic structure formed in the heat affected zone of welds clearly depends on the
heat cycle each point in the weld has been through, the type of material, and the initial structure.
For the same material and welding technique the major variations encountered relate to the
distance from the weld zone as the heat cycles are reproduced in the same way for points the
same distance from the weld bead.
Experiments have, however, shown that generally the metal undergoes structural transformations
up to a distance from the weld line that does not normally exceed 3 mm. The HAZ can extend to 7
– 8 mm in particular cases, such as when high heat procedures are used.
In the case of carbon steels and low alloy steels the most problematic phenomenon found in both
macro and micrographic analysis is the change in the size of the grains in relation to the distance
from the weld line.
Bearing in mind therefore that the width of the HAZ depends on the welding procedure and
parameters, let’s look at what happens in a C-Mn steel, with a minimum breaking strength of 490
N/mm2.
Let’s imagine that the transition zone is broken down into 3 layers, each about 1 mm thick (Fig. 74).
49
Fig. 74
The first layer, adjacent to the weld zone is said to be “overheated”. In fact all its points are taken
to very high temperature cycles resulting in the austenitic grain becoming larger (Fig. 75).
Fig. 75
As one moves away from the weld zone, this overheating decreases as does the coarseness of the
grain.
In this area the mechanical characteristics are inferior to those in the basic material.
The second or middle layer is said to be “normalised” (Fig. 76).
Its points are taken to temperatures that exceed the A3 line, but not by much. This means that the
heat cycles more or less reproduce those found in the various phases of normalisation heat
treatment.
The final structure of this second layer is fine and granular, with mechanical characteristics that are
better than the previous layer, and in some cases even better than those of the basic metal.
Fusion line
Fusion zone
Coarse grain zone
Fine grain recrystallised zone
Partially austenitised zone
C(%)
T(°C)
Liquidus
Liquidus + γγγγ
γγγγ
γ γ γ γ + Fe3C
α α α α + Fe3C
αααα+ γ γ γ γ
T max during thermal cycle
HAZ
Coarse grain zone
The extension of the coarse grain zone dipends on the holding time between 1100°C and 1500°C.
50
Fig. 76
The third or outermost layer is said to be “partially austenitised” (Fig. 77).
The points in this layer reach temperatures that fall between lines A1 and A3.
Fig. 77
In this zone, globular pearlite is formed and the mechanical characteristics are generally superior to
those of the unaltered basic material that contains laminar pearlite.
Further away from the weld zone, no structural alteration is found in the basic material.
C(%)
T(°C)
Liquidus
Liquidus + γγγγ
γγγγ
γ γ γ γ + Fe3C
α α α α + Fe3C
αααα+ γ γ γ γ
T max during thermal cycle
HAZ
The extension of the fine grain recrystallised zonedepends on the holding time between 850°C and 1100° C.
Fine recrystallized zone
C(%)
Liquidus
Liquidus + γγγγ
γγγγ
γ γ γ γ + Fe3C
α α α α + Fe3C
αααα+ γ γ γ γ
T max during thermal cycle
HAZ
The extension of the partially austenitised zone dipends on the holding time between Ac1 and Ac3.
Partially austenitised zone
T(°C)
51
In the case of steels with a high content of carbon or alloying elements, with minimum breaking
strengths exceeding 490 N/mm2, the corresponding TTT curves are moved to the right compared to
those for mild steels so it is more probable that hardened or fragile structures will be found in the
HAZ once cooling has finished.
Briefly, the parameters that determine the structure of the HAZ are:
- heating speed: as the heating speed increases, the tendency of the grain to be smaller
increases since the time spent at higher temperatures is shorter. At the same time, the time
available for impurities to be dissolved is reduced and as these constitute the starting point for
pearlite transformation, this transformation is favoured and the hardening capacity is reduced.
- maximum temperature reached and time spent at high temperatures: both these factors act
as elements that tend to favour the grain being more coarse, therefore acting favourably in
terms of the hardening capacity of the structure.
- carbon content: hardness intervenes more easily in direct proportion to the carbon content. In
addition, if the heating speed is high during alpha and gamma transformation the carbon does
not have the time to distribute itself evenly in the austenite and the areas that are richer in
carbon favour hardening.
- speed of cooling: as is known, this acts in the same way as the carbon content.
The simplest way to estimate the heat changes experimentally consists of taking hardness readings
on items welded using the same procedure as that adopted for structural joints.
Excessive hardening (Vickers hardness of more than 300) is cause for alarm, and indicates that less
severe welding methods and conditions must be sought.
52
2. Welding Defects
Hydrogen in welding and cold cracking
As indicated in previous paragraphs in relation to oxygen and nitrogen, the solubility of hydrogen in
steel varies in relation to temperature and crystallographic shapes.
Its solubility will therefore be high in liquid steel, low in delta ferrite, greater in austenite, and almost
zero in alpha ferrite at ambient temperature (Fig. 78).
If the steel absorbs a significant quantity of hydrogen in its liquid state and then, as happens in
welding, it solidifies and cools very quickly, part of the hydrogen absorbed does not leave the
solution quickly enough and is trapped in an over-saturated state in the solidified steel.
The hydrogen in the over-saturated solution causes a distortion in the steel’s crystalline lattice. This
distortion does not cause any problems in austenite as it is ductile and has space in its lattice that
can accommodate the hydrogen atoms without excessive distortion.
In ferrite however, and even more so in martensite that is already fragile and distorted, hydrogen
causes a significant increase in fragility of the structure, which may result in flaws forming as we will
see later on.
We also know that hydrogen atoms are very small and it is therefore possible for them to spread
through the metal matrix, moving towards the surface of the weld or to micro or macro cavities in
the weld itself.
The diffusion of hydrogen also depends on the temperature and the structure of the steel.
This phenomenon increases as temperature rises and, at the same temperature, is much more
widespread in ferrite than in austenite, where it is reduced to almost zero already at 400°C. In
ferrite, hydrogen is able to spread even at ambient temperature (Fig. 79).
Fig. 78 Fig. 79
It is therefore clear that, if austenite and ferrite are both present at a given temperature, the
hydrogen will tend to spread out from the ferrite into the austenite, both due to its greater capacity
for spreading in the ferrite and its greater solubility in austenite.
If one wishes to eliminate over-saturated hydrogen from a weld without waiting for it to slowly
diffuse to the outside naturally at ambient temperature, the temperature need simply be
increased, which is exactly what is done in heat treatment that eliminates the so-called diffusible
hydrogen.
This treatment involves maintaining a temperature of 250°C for a long time, the period varying
between 6 and 16 hours.
0
4
8
12
16
20
24
-200 0
200
400
600
800
1000
1200
°C
CCC
CFC
53
“Diffusible hydrogen” is the hydrogen that spontaneously escapes from the weld over a long time
at ambient temperature. Hydrogen is responsible for the weld defects that we will be looking at
next.
The sum of the diffusible hydrogen and the residual hydrogen is referred to as “total hydrogen”.
Residual hydrogen is the hydrogen that leaves the solid material after treatment at 650°C, which
may be carried out after all the diffusible hydrogen has escaped.
Residual hydrogen is the hydrogen that remains in the cavities and inclusions in the weld at
ambient temperature.
Total hydrogen does not include a certain quantity of “fixed hydrogen”, which could only be
removed by melting steel in a vacuum.
Hydrogen absorption in welding
a) Oxygen-acetylene Welding
In this welding procedure the acetylene (C2H2) reacts with the oxygen (O2) and produces carbon
monoxide and hydrogen, which is easily absorbed by the welding puddle. Due to the slow cooling
speed and the high diffusivity of hydrogen at high temperatures, the hydrogen has sufficient time
available to escape from the molten puddle and so hydrogen rarely creates any problems in
oxygen-acetylene welding.
b) Welding with coated electrodes
In this case hydrogen is released by dampness in the coating that may contain mineral
crystallisation water, or ambient dampness in the place in which the electrodes are kept.
The speed of solidification is rather high and so only a small amount of the hydrogen is able to
escape to the outside while the weld is cooling.
The only “low hydrogen” electrodes are those with basic coatings, while all the others contain high
quantities of hydrogen. In the case of basic electrodes it is essential that the coating be properly
dry and, as they are highly hydroscopic, they must be kept in a dry place and dried before use if
necessary (Fig. 80).
Fig. 80
c) Submerged arc welding
As is the case for electrodes, in this case the hydrogen also comes from dampness in the flow.
In the case of both electrodes and flows for a submerged arc there is a procedure that is used to
determine the amount of dampness in the flow itself, known as a total water test.
The dampness values allowed for coated electrodes vary between 0,15% and 0,2%, and from 0,07%
and 0,15% for the flows.
As the reader will have noticed, lower values are prescribed for submerged arc welds and this is
because the part of the flow that melts in this procedure is higher in percentage terms than the
part that melts when using coated electrodes.
d) Welding under a gas barrier
In this case the hydrogen may be drawn from the humidity in the gas or, when welding with a flux-
cored wire, by the humidity in the flow.
54
Cooling takes place quickly and the comments made for coated electrodes apply here as well.
It is therefore essential to use protective atmospheres with low humidity percentages, and in
practice gasses are used with a dew point of less than – 40°C at least.
Once welding has been completed and using various procedures, the hydrogen content left in
solidified alpha iron, expressed in cubic centimetres per 100 grams of metal deposited, can be
determined using conventional harmonised tests.
In terms of the values found, welding materials are classified as follows:
- < 5 - very low hydrogen
- 5 to 15 - low hydrogen
- > 15 - high hydrogen
Hydrogen problems in welding
Blowholes
Blowholes are not a type of defect that is
peculiar to hydrogen insofar as they can also
be caused by carbon monoxide and carbon
dioxide for example.
In the case of hydrogen the gaseous inclusions
originate from hydrogen sulphide (H2S) or
methane (CH4) and may cause elongated
blowholes commonly known as “worm holes”
(Fig. 81).
Fig. 81
Flakes
We have seen how over-saturation hydrogen in welding tends to diffuse during cooling. This
diffusion not only moves towards the outside but also passes through the metal matrix towards the
cavities, micro cavities, or micro inclusions that are always to be found in welds.
The atomic hydrogen that gets into these cavities changes to a stable molecular state at low
temperatures and tends to increase pressure to the extent that even very high pressures may be
reached. This strong pressure acts spherically in all directions, creating a loading that acts on three
axes, making the steel more fragile and prone to local failure due to de-cohesion.
When plastic deformation takes place, such
as in a bending or tensile test (Fig. 82), the
areas around the cavity may therefore break
prematurely in a very fragile manner.
After such a break, a trace of the micro
cavity that housed the hydrogen can be
seen on the broken section of the test piece,
in the form of one or more light coloured
spots with a diameter of 1-2 mm and a black
dot in the centre.
This defect is known as a “flake”. The light
coloured part of the flake is nothing other
than the appearance of a localised fragile
break without deformation that stands out
against the break section around it that broke
with plastic deformation and appears darker
and more fibrous.
Fig. 82
The presence of flakes does not significantly alter yield strength or tensile strength, but it does cause
a reduction in elongation and necking down, which is very understandable when one considers
that some areas broke due to de-cohesion and therefore played no part in plastic deformation.
It should be pointed out that flakes are only formed during plastic deformation and so they are not
a great defect in welded structures.
55
In fact, a structure is rarely put into use before the diffusible hydrogen has had the time to escape
from the weld (remember the heat treatment for releasing diffusible hydrogen) and operating
stresses almost never cause plastic deformation. Clearly it is only in exceptional cases that flakes
pose any danger.
However, this defect represents a circumstance that points to the presence of hydrogen with all
the problems this may create, and so it must be taken into account when carrying out welding
procedure approval tests.
Micro cracks in the weld zone
When significant quantities of diffusible hydrogen are involved, such as for example in the deposit
made when using electrodes with a rutile or cellulose coating, if cooling takes place quickly the
hydrogen that is rapidly accumulated in the spaces and around these spaces in the ferrite matrix
can cause the formation of small discontinuities when subjected to longitudinal shrinkage forces.
This phenomenon is more marked when the atmosphere is less protected against nitrogen. In fact,
the presence of nitrogen in the weld zone promotes this phenomenon.
These defects are known as hydrogen micro cracks and are found in the centre of each pass and
laid out transversely to the weld (Fig. 83).
Due to their orientation and small size, they
have a limited influence on the ductility
determined in tensile or bending tests carried
out transversely to the weld. In longitudinal
tensile and bending tests, however, the micro
cracks link up causing macroscopic breaks.
In addition, they reduce the steel’s resistance
to fatigue as they provide multiple starting
points for breakage and also reduce
resilience, when found in large numbers.
In general, the influence of micro cracks is
very limited to the point that no breaks are
found in practice that can be attributed to
this defect.
Fig. 83
Cracks in the HAZ (cold cracks)
Cracks in the HAZ when either cold or hot, which we will look at next, are defects of a metallurgical
nature and, as such, independent of the operator.
These are two-dimensional defects and so, since they represent a reduction in the resistance
section and provide a starting point for breaking, they are always unacceptable.
When a material with a carbon or alloying element content is such that, with a rather severe
welding heat cycle, it forms a hardened, fragile structure in the HAZ, cracks may form in the HAZ
itself in the immediate proximity of the weld zone.
These cracks are more likely to form the harder the HAZ is and the higher its martensite content and
hydrogen content are; they are known either as hydrogen cracks, due to their location under the
bead or, more frequently, cold cracks due to the low temperature at which they form.
Materials that manifest phenomena of hardening are particularly prone to these cracks, and so this
includes materials that undergo allotropic transformations in the solid state.
To explain the mechanism behind the cold cracks, it must be borne in mind first of all that hydrogen
is more soluble in austenite structures than in ferrite structures, although its diffusion capacity is
greater in the latter. In addition, for the same structure, the diffusion capability of hydrogen
increases as temperature increases.
In the paragraph dealing with martensitic structures we dealt with hardening factors and especially
carbon content, which contribute to moving the TTT curves to the right, for continuous cooling of a
certain material.
Normally the filler material has a lower carbon content than the basic material and so the TTT
curves for the WZ shift to the left (that is, towards shorter times) and upwards (towards higher
temperatures) compared to the curves for the basic material (Fig. 84).
56
Given their proximity, the temperatures at a
given moment in the WZ and the adjacent
HAZ are assumed to be equal, in the
immediate proximity of the line separating the
two zones, and therefore the cooling curves
will also be identical.
The weld zone for which the TTT curves have
been moved to the left undergoes
transformation from austenite to ferrite before
the HAZ, and if the cooling speed is
sufficiently high the entire weld zone
undergoes ferrite pearlite transformation while
the heat affected zone becomes completely
martensitic.
Fig. 84
Let’s see what happens following these transformations (Fig. 85).
The weld zone that was initially rich in hydrogen is transformed from austenite to ferrite first and
tends to expel the hydrogen due to its reduced solubility and the simultaneous increase in diffusion
capability.
The HAZ is mainly austenite and initially poor in hydrogen and so readily absorbs that coming from
the WZ. Due to the high solubility in austenite and therefore low diffusion capacity, the hydrogen
tends to remain concentrated in the immediate vicinity of the weld zone.
When the austenite becomes martensitic,
which occurs instantaneously along with an
increase in volume, the hydrogen is trapped
in the lattice in an over-saturated solution and
tends to escape very slowly. This escape
occurs more slowly where the martensite is
formed at lower temperatures as its diffusion
capacity decreases with temperature.
The HAZ is therefore mainly made up of
martensite with a hard and fragile structure
that is made even more fragile by the
presence of hydrogen in an over-saturated
solution, subject to both the residual tensions
of welding and those related to the increase
in volume.
Fig. 85
Under these conditions breaks may well develop that arise precisely where the highest
concentration of hydrogen is to be found, namely under the bead (Fig. 86 & 87).
FZ
HAZ
log(t)
T
αααα γγγγ
αααα γγγγ H
α+γα+γα+γα+γ
57
Fig. 86
Fig. 87
Cold cracks may also be found in the weld zone and this may occur when the weld is done using
filler material that is identical to the basic material.
In this case the TTT curves coincide and the weld zone is transformed at the same time as the HAZ
or even after, not allowing the hydrogen to spread into the basic material.
If martensitic transformation occurs, cold cracks arise wherever the hydrogen is located, i.e. in the
weld zone.
From the above it is clear that three factors are required for cracks to form in the heat affected
zone: tension, hydrogen, and hardness.
Tension
No efficacious remedies are available for tensile forces, as it has been shown experimentally that
the tensile forces involved in the binding in the weld and those related to the increase in volume in
the changeover from austenite to martensite are sufficient to cause cold cracks.
Any precautions when welding for free contraction do not suffice to avoid the danger of breaks
occurring.
Hyrogen
The hydrogen that migrates from the weld
zone to the heat affected zone is nothing
other than hydrogen in the filler material that
can be diffused.
Therefore, to obtain a low presence of
hydrogen, steps must be taken as indicated
beforehand, namely fully dry basic flows and
electrodes, protective gasses with a low dew
point, etc.
It is also good practice to use preheating,
which, as we already know, facilitates the
diffusion of hydrogen to the outside.
Fig. 88
Hardness
The harder the martensitic structure is, the more likely it is that cracks will occur, and it has been
shown that this hardness is directly related to the carbon content but is independent of the content
of other alloying elements.
The martensite content depends on hardening factors that have been dealt with before. We said
that the action of these factors takes the form of moving the TTT curves to the right and lowering
the interval in which the hardness structures are formed (Ms – Mf).
This interval is very important as it demarcates the temperature range over which martensite forms
in the sense that all the martensite that we find after a weld has formed for example, has formed
below a certain temperature (Ms) and above another temperature (Mf).
Reducing the Ms – Mf temperature difference and the Mf temperature in particular results in the
formation of martensite at a very low temperature and therefore when the diffusivity of hydrogen is
58
almost non-existent; the hydrogen is concentrated in the last remaining austenite areas where it is
trapped in a highly over-saturated solution in the last martensite parts formed due to transformation
of the austenite areas referred to above.
Bearing in mind that the element that has the greatest effect in lowering the range over which
martensite is formed is carbon, one realises how much this has a negative influence. It moves the
TTT curves to the right, moves the temperature interval Ms – Mf down, and increases the hardness of
the martensite.
Naturally, the rate of cooling plays a very important role in the transformation of martensite, along
with all the comments we covered when dealing with heat cycles.
It should be noted that the cooling speed
between A3 and Ms determines the amount
of martensite that forms in the HAZ, while the
rate of cooling in the transformation interval
determines the speed with which this
transformation takes place.
This is an important factor because at a high
cooling speed the amount of hydrogen
trapped in the hardened structure is greater.
In this case one should not lose sight of the
beneficial effect of preheating, which not
only slows cooling at high temperatures, but
does the same at low temperatures as well,
including the transformation range (Fig. 89).
Fig. 89
Hot Cracks
As with cold cracks, hot cracks are also of a metallurgical nature, two-dimensional and
unacceptable.
Unlike cold cracks they can affect all materials and they form, as indicated by the name, during
the initial phases of cooling.
The ever-present residual stresses from welding and the solidification interval for the material cause
these cracks.
In fact, we know from studying state diagrams that while a pure material solidifies at an exact
temperature, an alloy solidifies over a temperature interval that varies depending on the
percentage content of the alloy components.
This interval is extended when the alloy includes impurities which, as we have seen previously,
cause dendritic segregation phenomena.
The increase in the solidification interval is
therefore made up of a liquid veil of impurities
that one finds at the edges of the grain when
solidification has been completed which, in
the case of sulphur, only occurs at 988°C.
The liquid veil obviously reduces the resistant
section of the joint, reducing cohesion
between grains.
If shrinkage is prevented, as always happens
to a greater or lesser extent in welds, the
residual tensile forces act in the gaps
between the grains already in contact
making up solid bridges (Fig. 90).
Fig. 90
59
If their section is similar and remains so for too
long (that is, if the residual liquid does not
solidify until a significantly lower temperature
is reached, when the tensile forces due to
shrinkage are high) or if these solid bridges do
not provide the ductility necessary to
compensate for the gradual increase in the
tensile forces by means of plastic
deformation, the material may not be able to
withstand these forces and the grains may
become detached, forming a crack (Fig. 91
& 92).
Fig. 91
Fig. 92 – examples of hot cracks
We have referred before to compounds that melt at low temperatures formed by sulphur; in
general these elements are eutectic, namely ferrous monosulphide (FeS) that segregates
preferably in a film shape around the grains being formed.
The higher the sulphur content of the weld puddle, the greater the extent of this liquid film and,
consequently, the more slender the solid bridges will be.
Phosphorous has a lesser influence over this phenomenon than sulphur as the low-melting
compounds it forms with iron (FeP) have a higher solidification temperature.
Carbon also has a harmful effect as at high temperatures it causes an increase in strength and a
reduction in ductility. This results in a reduction in the capacity to assist with the shrinkage by means
of plastic elongation of the solid bridges between the grains.
Thus the simultaneous presence of relatively high percentages of sulphur and carbon is particularly
harmful.
Another compound that has a similar effect to carbon is silicon as it significantly increases fragility.
The grain size also plays an important role in hot cracks. In fact, the larger the grain the smaller the
contact surface and thus the lower the cohesion between grains, since in finer structures the solid
bridges would be more numerous and better distributed.
60
This means that, in the case of large grains, the shrinkage forces act on a smaller number of solid
bridges that are more poorly distributed, with a clearly greater danger of flaws.
Another parameter that must be taken into
account is the section of the joint, which is
defined as the shape factor, i.e. the ratio
between the width of the bead and its height
W/D (Fig. 93).
If the joint is too narrow, cracks may occur on
the contact surface of the columnar
dendrites that are solidifying due to the
different position the dendrites are forced to
take up.
The shape factor begins to be critical for hot
cracks when it has a value of less than 1.
The shape factor is influenced not only by the
preparation of the caulker but also by the
welding speed (Fig. 94).
Fig. 93
Fig. 94
As to the position of hot cracks, it should be noted that while cold cracks are typical of the HAZ, hot
cracks are almost always found in the WZ.
Depending on their position in relation to the weld axis, these may be:
- longitudinal in relation to the root of the weld: these are typically found in T-joints in which
angular shrinkage is very low, and in butt-joints, especially if the first pass is too thin (Fig. 91 & 92).
- longitudinal along the axis in the direction of the weld: these are found in T- and butt-joints and
indicate low ductility in the area that solidifies last and that, being at high temperature when
the surrounding areas have already become rigid, are subject to the greatest plastic
deformation (Fig. 91-92, 95).
- inter-dendritic (also known as liquation cracks): these are the smallest and form between the
dendrites, generally not breaking the surface of the weld. These are the most frequent and
most insidious as they are difficult to detect. In fact they can only be found in destructive tests
since, due to their small size, they cannot be picked up in x-ray tests and ultrasound tests only
sometimes manage to detect their presence (Fig. 96).
- transverse: these generally develop in high tensile steels due to the longitudinal shrinkage
tensile forces. They often break the surface and generally only affect one surface layer (Fig.
91).
- crater: craters form in the ends of welds and normally these radiate out from the centre of the
crater, i.e. the area that solidifies last, thereby containing the greatest concentration of
impurities and subject to the greatest plastic deformation. In arc welding using coated
electrodes they may be due to the failure to protect the puddle and the concave shape it
takes up due to the electric arc being detached too quickly (Fig. 97).
All these cracks come in various sizes in the sense that you can have a microscopic crack like a
splinter that goes right through the entire joint.
In particular cases hot cracks may also occur in the heat affected area.
This happens most often when welding hardened and tempered stainless steel with a high carbon
content, but are rare in mild steels.
These are known as liquation cracks and consist, during welding, of the fusion of low-melting
impurities in the HAZ. This results in liquid films that cause a lack of solid continuity between the
grains in the HAZ and favour the formation of cracks due to residual tensile stress (Fig. 96).
61
Fig. 95 Fig. 96
Fig. 97
Remedies for avoiding hot cracks
The welding technique used can be of assistance in avoiding bonds that are too rigid and using a
welding sequence that allows welding with free shrinkage as far as possible.
Stress relieving treatment does not solve this problem as the cracks form at temperatures above
those used for this type of treatment.
With regard to low-melting impurities, as far as possible killed steels must be used as these have a
low S and P content, and effervescent steels must be avoided.
It is always advisable to use basic electrodes and flows that are slag forming and draw the
impurities into the slag, removing them from the weld puddle.
The specific heat build-up is also important as this determines the dilution and the final grain size.
We know that, when welding, a filler material that is purer than the basic material must be used
and so dilution must stay at low levels to avoid the impurities getting into the weld puddle as far as
possible.
The higher the heat build-up the slower solidification occurs, the larger the grain size ends up being
once cooling has been completed, the smaller the extent of the joints between the grains along
which the impurities segregate, and thus the higher the concentration of these and the probability
of hot cracks occurring.
The choice of the caulker is also important for hot cracks.
For example, in the case of a very thick steel plate in which the impurities are concentrated at the
centre of the thickness due to the segregation phenomena that occurred when the ingot or slab
62
from which it is made was solidifying, if the caulker used is prepared with X symmetrical, the first
pass brings with it the greatest heat build-up and thus greatest expansion and will be found right in
the centre of the band of impurities, creating a very impure weld puddle that has to withstand the
residual tensile stresses on its own.
If the piece is prepared with X dissymmetric, obviously the first pass will be in a purer zone, reducing
the risk of hot cracks forming (Fig. 98).
Fig. 98
Laminar Tearing
Laminar tearing are cracks normally found in laminated products, which occur in restricted joints
when the material is subjected to tensile forces perpendicular to the lamination plane, i.e. across
the thickness of the plate (commonly referred to as the “short transverse”) (Fig. 99).
This defect is typical of basic materials and occurs in the HAZ or in the nearby basic material.
The characteristic layout of the grains, more or less parallel to the rolling surfaces, distinguishes this
defect from cold cracks under the bead (Fig. 100).
Fig. 99 Fig. 100
The principal factors that tend to lead this defect arising are:
- the geometry of the joint, where the area around the weld zone is practically parallel to the
laminate’s surface, which means that the shrinkage tensile forces act in the direction of the
short transverse.
- material with great anisotropy in relation to its deformation characteristics. In this case the
material’s ductility in relation to its thickness may be insufficient to support the concentration of
tensile forces and deformations.
- localised tensile forces of significant size, developed by strongly bound welded joints following
shrinkage while cooling down. These take on greater importance when the structural member
is highly rigid and the weld zone has an extensive volume.
The tear normally begins in the plate near the well in the area around the inclusions laid out in the
plate parallel to the external surface.
In general the non-metallic inclusions that are most susceptible are manganese silicates and
aluminium, due to the deoxidisation process used on the steel and elongation during the
subsequent rolling.
63
In some cases laminar tearing starts in the heat affected zone due to the effect hydrogen has of
making the area more fragile and/or the formation of cold cracks.
These defects significantly reduce ductility in the short transverse direction, and when the material is
stressed in this direction the first breaks occur due to de-cohesion right on the inclusions which,
being closer to the weld, have to support the greater shrinkage tensile forces.
Subsequently, plastic deformation bands are concentrated in the areas stressed along three axes
at the apex of each micro cavity, producing a coalescence of the first breaks that are substantially
perpendicular to the inclusions (cutting action).
The three joints most often affected by this phenomenon, in descending order of criticality, are:
- full-penetration cross-joints and T-joints with 1/2V or K preparation
- full-penetration L-joints
- T-joints or cross-joints with fillet welds
The main steps that can be taken to reduce the danger of laminar tearing are:
- the use of steel with improved characteristics across the thickness (low sulphur content, globular
inclusions, necking down in tensile test pieces over the thickness exceeding 25 – 35%)
- the use of pieces with a special geometry suitably shaped for critical joints (Fig. 101)
- reducing the size of the weld zone (Fig. 101)
- the use of filler materials with greater ductility and lower yield points than the basic material
- the use of a suitable order of execution of the passes in order to reduce the local stresses that
are applied to layers of the weld zone that are more ductile than the basic material (Fig. 102)
- reducing shrinkage tensile forces by using the “buttering” technique, which involves depositing
a layer of filler material with a low yield point and high ductility on the surface of the piece
before the actual weld is formed (Fig. 103).
Fig. 101
Fig. 102
Fig. 103
64
3. Heat Phenomena in Welds (Shrinkage, Internal Tensile Forces) Shrinkage in Welds As is well known, the increase in temperature in a body is accompanied by an increase in its size,
while for any reduction in temperature the body contracts until it returns to its original size. In
welding, shrinkage is the change in size that a material undergoes as it goes from melting
temperature to ambient temperature.
The increase in the temperature of a piece of metal causes it to expand, and this may take place
uniformly only when the piece is heated up uniformly and there is nothing to stop its free expansion.
Where this is not the case, other phenomena come into play, which we will look at later.
Let’s look at a bar of steel that is unhindered and subject to uniform heating up. It increases in
length but when it cools down again it returns to its original size (Fig. 104 a).
A second piece secured at its two ends is put into a state of compression due to its expansion
being prevented. This compressive force increases as the temperature increases until the yield
point is reached.
Starting from this point, the bar undergoes irreversible plastic deformation when subjected to any
further increase in temperature, that is, it is upset (Fig. 104 b).
When the bar cools down it shrinks but its length ends up shorter than it was originally while the
diameter is greater. Once the bar has cooled down, it is free of any stress.
Finally, let’s look at a bar that is fixed at its ends in such a way that contraction is also prevented.
It is upset as before during heating, and when it is then cooled since it cannot shorten it is put into
tension, which is known as residual tension or internal tension (Fig. 104 c).
Fig. 104
Expansion at T max.
a
c
b
Final contraction
No residual deformation
No residual tension
Contraction residual deformation
No residual tension
Blocked residual deformation
Traction residual tension
σS
0
A B
C
T
traction
compression
65
When a piece of metal is heated up in a non-uniform manner its expansion is always restricted. In
fact, the heated part tends to expand more quickly than the surrounding cool parts, but it cannot
do so freely as it is connected to the rest of the piece.
At this point, if the temperature is raised enough the hotter part is upset and once it cools down its
fibres are tensioned, often at values that are very near its cold yield point.
Similar phenomena occur in welds since in this case too the heat build-up is localised and thus
heating up of the piece is highly non-uniform (Fig. 105).
Fig. 105
To simplify the study of heat phenomena that occur in welds and the effect they have on shrinkage
in welded joints, the shrinkage is broken down into three directions identified as follows:
- transverse shrinkage that acts across the welded joint on the plane of the plate
- longitudinal shrinkage that acts in the direction of the axis of the weld
- perpendicular shrinkage that acts in the direction of the thickness of the plate.
Only the first two are important in terms of the extent to which they occur and thus for the
consequences they cause. Let’s describe them properly (Fig. 106-107).
Fig. 106 Fig. 107
Transverse shrinkage
Two aspects can be looked at in studying this:
a) Transverse shrinkage per se
Let’s look at the phenomenon of transverse shrinkage resulting from a butt weld joining two plates
with straight edges. The movement of the heat source takes place longitudinally and at a constant
Welding bead
Joint axis
50 °C
1500 °C
700 °C
V •
HAZ
66
speed. The joint is formed in a single pass and the plates are not restricted in any way from the
outside or by tack welding (Fig. 108).
Fig. 108
During the welding process two zones can be identified on the plates in relation to the isotherm
ellipses bounded by the end of the minor axes. One (ABCE) is cooling and the other (AECF) is
heating up due to the heat source moving forward.
In the first zone there is a portion of the joint with t > 600°C (DE), with a yield point that is very low.
Looking at a strip of material in the DE area, by analogy with the restricted bar one can say that in
this area hot upsetting will take place due to the restriction caused by the surrounding material
(outside the 600°C isotherm), which will impede expansion and act as an anchor for the upsetting
due to further elongation of the side pieces that are being heated up.
When T decreases as the heat sources advances, the elastic limit of the strip tends to return to
normality, resulting in the strip contracting.
This contraction is strongest in the DE stretch on the source side, where the material at high T is
easily upset, while zone B, which is already cooling further, acts as a fulcrum, causing the plates
that are still free to move towards one another (scissor effect).
Transverse shrinkage is directly proportional to the width of the weld and the quantity of heat built
up, and is inversely proportional to the thickness of the weld.
b) Angular shrinkage
When looking at a single pass weld in section, one notes how the width of the bead varies along
the thickness of the joint, reaching its maximum width at the surface (Fig. 109). By considering
dividing the plate into parallel strata, due to the fact that transverse shrinkage is proportional to the
width of the bead one sees that the upper layers shrink more than the lower layers, causing angular
shrinkage.
Fig. 109
In fillet welds shrinkage takes the form of angular shrinkage (Fig. 110).
67
Fig. 110
The rotation of the plates (tendency to close) increases as the cross-section of the bead increases
and as the number of passes increases, since on the faces of the two pieces there are two zones
resulting from the heat cycles in relation to the thickness, one near the weld that will undergo hot
upsetting and the other further away that will simply deform in the elastic field.
Transverse shrinkage increases as the section of the weld increases, but much less quickly as the
thickness to be welded increases.
Longitudinal shrinkage
This always takes place under self-restricting conditions and is therefore generally the smallest of the
lot.
In fact, while in the case of transverse shrinkage the contraction of the material takes place to a
greater extent the less it is subject to outside restrictions (such as spot welds), in the case of
longitudinal shrinkage there is always a restriction imposed by the joint itself.
In this case the bead and the surrounding hot part tend to deform as shown in Fig. 111, but this
contraction is restricted by the surrounding cold part.
Shrinkage is minimal in this case and simultaneous and so very high residual tension results. For a
given section of weld bead, the extent of shrinkage decreases as the section of the plate increases
and, for the same cross-sectional area of plate, it increases as the area of the weld section
increases.
Fig. 111
Origin of residual tensions in welds
The restriction in a weld caused by un-uniform heating and cooling and the impossibility of
shrinkage developing unhindered generate tensions of a greater or lesser entity in the joint itself
that are known as residual tension or shrinkage forces.
Shrinkage of metal objects should take place in all directions but this is impeded to a greater or
lesser extent by external and internal restrictions, which gives rise to a complex state of stress in a
direction that is longitudinal, transverse, and perpendicular to the thickness.
68
Longitudinal tension
Longitudinal tension is the highest due to the severe self-restriction effect.
These are tensile forces in the welded joint and immediate vicinity and, due to logic of seeking
equilibrium, are in compression in the outermost areas (Fig. 112).
This type of tension remains constant for all sections parallel to the joint and drops to 0 at the ends.
For the same section of the weld zone, as the plate thickness increases and for the same plate
thickness, this tension decreases as the weld zone gets bigger.
Fig. 112
Transverse tension
Transverse tension is made up of tensile forces in the centre of the joint, becoming compression in
the external areas (Fig. 113).
Transverse tension is generally lower than longitudinal tension and largely depends on the degree
of external restriction.
In the case of unrestricted plates the tension decreases as the weld zone increases in size, and
therefore the specific heat build-up increases. In the case of restricted plates the opposite occurs.
Fig. 113
69
Effects of shrinkage and residual tension
When the outside loads imposed by the working conditions act on a welded structure they are
superimposed on the welding tension already in the joints.
If the working loads counteract the welding tension the latter is reduced and may even be nullified.
This is the most favourable case and makes use of the existing pre-tensioning.
In the case where the working loads are tensile like the welding tension, these tensile forces are
added to one another and may exceed the material’s yield point, causing plastic deformation.
For a better understanding of the influence of internal tension on strength of welded structures let’s
look at the following comparison (Fig. 114).
Fig. 114
Lets consider the longitudinal tension in a butt-joint which, as we have said before, is highest and in
tension at the centre of the bead and in compression at the edges.
This can be likened to a frame with two rigid beams A - B and three equal columns C – X – C1, of
which the central column X passes through the beams and can be tensioned using a nut up to its
yield point.
In this case the side columns C and C1 will be in compression.
By applying a load to beam A the central column X begins to elongate and yields plastically but
does not play any part, having already yielded, in supporting the load. One notes, however, that
the side columns are subject to greater compression to make up for the central column not making
any contribution. The following example is provided to further clarify the ideas involved:
- section of each column = 100 mm2
- yield point Rs = 240 N/mm2
The applied load will be 100 x 3 x 240 = 72000 N.
If the central column is strained to its yield point a pre-compression of -120 N/mm2 will be applied to
each of the side columns; this means that each side column may undergo an elastic change in size
equivalent to -120 to +240 N/mm2 and the load that can be supported by columns C and C1 will
be 100 x 2 x 360 = 72000 N.
This shows that the strength of the frame remains unaltered.
In some cases residual tension must be taken into account as it may pose a danger, and this is
normally done for structures:
- subject to fatigue
- subject to fragile breaking
- subject to stress corrosion
- subject to machining with precision machines
In these cases particular precautions must be taken or procedures adopted in order to obtain
welded joints with significantly low residual tension.
70
Practical methods for attenuating tension
Any remedies to avoid the consequences of residual tension must be broken down into three
operating phases:
Precautions before welding
The precautions a technician can take before welding a joint must meet the basic requirement of
“providing for shrinkage”, that is, they must create conditions that allow shrinkage to take place as
freely as possible, with the welded piece ending up in the exact position required.
Let’s take a quick look at the precautions most commonly used:
a) Prior deformation to allow unrestricted shrinkage
The pieces to be welded are given an equal
and opposite set-up or deformation to that
caused by the pieces shrinking after welding.
This simple arrangement makes it possible to
eliminate the effects of angular shrinkage in
butt-joints and fillet welds, and some
examples are shown in Fig. 115. The deviation
angle is normally just a few degrees and is
determined in each case depending on
operating procedures and conditions. In
general this is based on experience gained in
previous cases.
Fig. 115
b) Creating an elastic zone near the joint
In this case, in order to allow for shrinkage an elastic zone can be created in the element that
deforms more easily thereby absorbing the shrinkage tensile forces. Examples of this procedure
are shown in Fig. 116 a, (where a slight undulation has been created in the thinner plate), in Fig. 116
b/c (where a lipped joint is formed rather than a T-joint, or with upturns at the ends of the plates
rather than straight edges meeting in a butt-joint, guaranteeing significant elasticity), in Fig. 117
(where, when welding pipes to piping plates a groove is formed to allow for shrinkage), and finally
in Fig. 31 (where, when connecting a circular element to a large plate it is slightly domed so that
the weld around the periphery tends to flatten it out).
Fig. 116
Fig. 117
71
Fig. 118
c) Useful arrangement of the weld bead
A typical example of how positioning the
weld bead differently can overcome the
problems associated with shrinkage is shown
in Fig. 119. In this case moving the bead from
position A to position B allows angular
shrinkage to be avoided and, if the
connection diameter is enlarged slightly on
both pipes, throttling due to longitudinal
shrinkage is avoided.
Fig. 119
d) Breaking down into panels
In large complex welded structures such as shipbuilding, more simple parts are welded in which
deformation can be limited more easily. This involves forming panels or finished blocks. Once these
panels have been formed, it is easier to control and allow for all the shrinkage that occurs, in the
overall structure.
Precautions during welding
a) Choosing the procedure
Angular shrinkage largely depends on the welding procedure and operating method used. Where
possible it is best to use welds that are symmetrical about the axis and on the plane that you do not
want to deform. For example, X-shaped preparations are recommended for butt-joints, especially
when dealing with thick plates. In the case of angled joints, angular shrinkage can be
compensated for by welding the two opposite corners simultaneously, thereby restricting the
angular deformation of the plate. In the case of cross-joints one need simply weld the two
opposite corners across the vertex alternately or simultaneously. The welding procedure also has a
great influence on the extent of shrinkage as we already know that when caulking is thinner there is
less transverse shrinkage directly related to the external restriction on the joint. While little can be
done in terms of the setting up of a joint in a restricted situation, in terms of procedure something
can be done when the pieces are free to move and therefore only longitudinal tensile forces come
into play due to self-restriction by the cool parts. In this case some benefit can be gained by
starting the weld from the centre and working outwards to the ends rather than going from one
end to the other. In any event, the important thing in any weld is that it must be made up of both
basic material and filler material that allow a certain degree of plastic deformation that is able to
absorb the shrinkage forces that are created, which are increased where the pieces are restricted.
High plasticity of the weld must in all cases be found not only when the work has been completed
but also during the welding itself as, for example, the tendency for flaws to be created is greater in
the first pass than in later passes (see the section on hot cracks).
b) Preparation of the joint and working sequence
The sequence in which work is done on the joint is important as it allow the maximum provision
possible to be made for shrinkage in a predetermined preferential direction. For example, when
working on a large flat panel as shown in Fig. 120, you can see how you should not start working on
welding the long edges without first having welded the transverse joint that finishes on these edges,
in order to allow for as much free shrinkage movement as possible on the transverse joint along its
axis. Another example is when welding an I beam (Fig. 121). When welding fragile materials,
72
localised preheating can also be used in order to reduce the welding tension by allowing for
shrinkage. This heating must not be terminated when the joint has reached the required
temperature, but must be kept up for the entire duration of the welding process. When repairing
cast iron pieces, for example, heating must be applied in order to increase the gap between the
edges to be welded. Once welding is completed and heating is terminated, the piece can shrink
uniformly and without any residual tension (Fig. 122).
Fig. 120
Fig. 121
Fig. 122
Precautions after welding
a) Shrinkage heat
Shrinkage heat involves localised heating that is used to shorten the fibres in the heated material
when it has cooled down, by means of hot upsetting. One typical example is the T-joint shown in
Fig. 123; the curvature taken up by the joint following shrinkage of the weld can be significantly
reduced to the point of straightening the piece altogether, by applying shrinkage heat along the
most elongated surface. At the points at which the material is heated the temperature reaches as
To be welded
Zone to be
preheated
73
high as 800°C, which allows expansion and therefore upsetting of the material, and reducing the
length during cooling (Fig. 124).
Fig. 123
Fig. 124
b) Expansion treatment in a furnace at 600-650°C
The principle behind expansion in a furnace is based on the fact that, as we already know, the
yield point for materials depends on temperature and decreases as temperature increases. In the
range of temperatures at which this heat treatment takes place the yield point is very low, of the
order of 40-50 N/mm2.
By putting the entire piece into the furnace (Fig. 125) and heating it to the temperatures indicated
above and keeping it at that temperature for a certain period of time (Fig. 126 & 127), the tension is
released and reduced to at least the yield point for the material at the temperature reached. The
fact that this tension can be released indicates that once expansion has been completed the item
can show plastic deformation, which means that the dimensions of the item can change to a
greater or lesser extent. One should also remember that during expansion heat treatment, the
benefits of tempering also occur, thereby eliminating the areas of hardness and improving the
material’s mechanical characteristics. The speed of heating up, the time the expansion
temperature is maintained, and the rate of cooling are all very important in terms of the success of
this treatment and are generally related to the thickness of the item.
Fig. 126 – source: RINA Rules
74
Fig. 125 Fig. 127 – source: RINA Rules
c) Treating local expansion in pipes
Localised expansion treatment is used when the entire piece cannot be given expansion treatment
in a furnace. It is a very particular method and is only used when it can be of real benefit to the
item. As one can imagine, if a plate is heated up locally to temperatures of 600 – 650°C, the hot
material expands and, if it is restricted from doing so by the colder parts, it is upset.
When the piece subsequently cools down, the tension will
not be limited to that already existing but will be added to
by the tension resulting from shrinkage of the upset material
opposed by the cold mass around it.
In practice therefore, the only case in which localised
expansion treatment is used is for circumferential joints in
pipes or for cylindrical structures that are free to expand
longitudinally (Fig. 128). Clearly, localised treatment is not
applied to longitudinal joints in pipes as the points made
relating to plates would apply in this case as well.
Fig. 128
d) Flame stretching for longitudinal tension only
This principle is based on the fact that applying a tensile load acting along the axis of the bead
causes plastic sliding of the stretched area which, once the load is released, results in the
longitudinal tension being redistributed, with higher tension being reduced. In the case of this
treatment the joint is stretched by heating the two strips on either side of the weld bead. This
causes them to expand due to the heat, resulting in plastic sliding in the central stretched area.
The heat is therefore only used to cause a mechanical pulling effect on the welded joint. Naturally,
given the problems that can arise in the case of localised heating it is best to keep these
temperatures below 200 - 250°C in order to avoid problems related to upsetting the material. This
treatment reduces longitudinal tension, while transverse tension remains unaltered.
e) Overloading or over-compression (mechanical stretching)
In this case the item is loaded or put into compression in order to reach loads that exceed the
working loads and to exceed the elastic limit at the points at which the greatest tension is
concentrated. At these critical points, since the tension due to overloading is added to the
residual weld tension, local plastic deformation occurs that may either significantly reduce the
residual tension involved or leave residual compression forces at the vertex of any notch that
75
reduces the danger posed by the same. Naturally this particular type of treatment must be carried
out on the basis of carefully designed procedures that also include the use of suitable checks (e.g.
ultrasound) in order to avoid problems due to existing defects in the structure.
f) Hammering the bead
Hammering can provide a double effect – it can straighten deformed pieces and reduce residual
tension, causing localised plastic deformation. This treatment must be carried out for each pass
and by skilled artisans to avoid problems related to creating tension in adjacent areas, strain
hardening, and ageing of the material. Due to the risks involved, hammering treatment is generally
not recommended (Fig. 129).
Fig. 129
76
4. Welding Processes
Electric Arc
Before going on to deal with various welding processes we must introduce the concept of the
electric arc, which is the physical phenomenon on which most welding processes in use today are
based. The electric arc is the visible manifestation of the passage of an electric current through an
ionised gas medium.
Let’s consider two refractory electrodes, for example made of carbon or tungsten, placed at a
certain distance one from the other and connected to a source of direct or alternating current.
When the circuit is open there is no current flow unless the applied tension reaches extremely high
values (thousands of volts/mm) high enough to rip electrons from their orbits around the nuclei of
the atoms making up the electrode.
If we heat the electrodes up, however, the force of attraction between nuclei and electrons is
reduced and once a certain temperature is reached, spontaneous emission of electrons is
achieved through the thermo-ionic effect.
Under these conditions it is enough to apply very low tension (in the order of tens of volts) to
accelerate free electrons away from the negative pole towards the positive one and obtain a flow
of current.
The air or gas molecules, struck by the fast electrons, become ionised giving rise to new electrons
and positive ions, that is fresh charged particles that will migrate towards the electrode of opposite
polarity. The collision between particles develops heat and an intense emission of light.
The large quantity of heat released confers continuity on the phenomenon, called “electric arc”
(Fig. 130).
Starting the electric arc is achieved by briefly short-circuiting the two electrodes. Through the
“Joule effect” the extremities in contact with one another generate fierce heat and by distancing
the electrodes the arc is triggered and remains active with a current flow of a few dozen volts (Fig.
131).
Starting the arc and keeping it alive means overcoming the dielectric rigidity of the air or the
resistance that the air itself opposes to the current flow.
Hence the ionising potential of gases or the thermo-ionic potential of certain materials is exploited.
Fig. 130
Fig. 131
77
The stability of the arc from a physical point of view is in fact influenced by two factors:
- Ionising energy of the gas
This is the energy that the gas requires to become ionised, i.e. the energy it takes to split the
molecules of the gas itself, and is related to the shielding gas typical of the welding process.
Clearly a gas with a lower ionising energy will make it easier to stabilise the electric arc.
In processes where a shielding gas is not used, such as when covered electrodes are being used,
the ability of the elements of the coating to become ionised is important.
Elements such as Na and K are employed because these have a high thermo-ionic potential and
are easily ionised and contribute to igniting the arc and keeping it stable.
- Temperature of the arc
If the temperature of the arc is high, then it will be more stable because the thermo-ionic potential,
or the ability of metals to emit electrons, is greater.
The temperature of the arc depends to a great extent on the thermal conductivity of the gas and
in fact less conductive gases correspond to higher temperatures.
To keep things simple we have illustrated the electric arc making reference to two identical
refractory electrodes whereas in manual and automatic arc welding we would find a fusible
electrode on one side and a base metal on the other.
Electric supply may be direct or alternating current; when direct current is used the polarity can be
direct (electrode connected to the negative pole) or inverse (electrode connected to the positive
pole). The atmosphere of an arc with a temperature of at least 5538°C is called plasma; more
properly said, plasma is the charge-carrying particles.
During experiments it was found that in an electric arc the plasma flow is anodic, that is it flows from
the anode (positive pole) to the cathode (negative pole) (Fig. 132).
Additionally, the cathode spot is always smaller than the anode one.
Fig. 132
These two factors tied to the geometry and physical state of the arc influence two of the electric
arc’s operative aspects: transfer of weld metal and penetration.
Transfer of the weld metal, following the flow of plasma, is for the most part anodic so in the case of
a fusible electrode it works better with inverse polarity.
Penetration is also greater in the case of inverse polarity because the cathode spot is smaller and
the energy is therefore more concentrated.
The Volt-Ampere characteristic of the arc is shown in Fig. 133 and represents the two values, tension
and current, that give rise to the arc. Here we will explain the phenomenon.
−−−−
+ + −−−−
anode
anode catode
catode
DCEN DCEP
78
Fig. 133 Fig. 134
We know that the resistance R of any conductor is expressed through the formula:
R = ρ L/S
where:
r = the specific resistance of the conductor
L = length
S = transversal section
In the case of the electric arc, as the current increases the transversal section S of the arc
progressively expands while, at the same rate, the temperature increases as does the percentage
of ionised particles with a consequential diminution of ρ.
The arc’s resistance R is therefore a diminishing function of the current and the diminution of R is so
important that even the product V = RI decreases as I increases.
The transversal expansion of the arc mentioned above tends however towards a limit that we can
define as “saturation” in correspondence with which complete ionisation of the gas takes place.
Beyond saturation the tension of the arc starts to increase in perfect agreement with the Law of
Ohm: V = RI.
The arc is said to be in free regime prior to saturation, after which the regime is throttled.
The position of this characteristic in the V-I diagram depends on the length of the arc, the
resistance of which increases as the length grows moving the characteristic point higher up in the
diagram.
With equal current, the tension is higher the longer the arc is.
Welding processes work on the ascending curve side of the diagram and therefore beyond
saturation point. As we have seen, the characteristic of an arc with a given length is a curve on
the V-I diagram the points of which represent all and only the voltage and current values at which
the arc can work. In an identical manner the V-I diagram curve that sets out the voltages and
current output of a generator is the generator’s external characteristic.
As a consequence, operating an arc off a given generator will be possible only in those points
where both the requirements of the arc and the generator’s performance are satisfied at the same
time, so only where the two curves intersect.
The static external characteristic of the generator is a negatively inclined line that joins the point
representing no-load voltage Uo to the one representing the short-circuit voltage Icc (Fig. 134).
Drawing the characteristic curves of both the generator and the arc on the same diagram, these
will intersect at two points of which only one (P2) is stable Fig. 135).
In fact if we suppose a positive increase is applied to the current so that P2 migrates to A on the
arc’s characteristic, the generator’s operative point migrates to B: the machine’s voltage is less
than that required by the arc so there is a diminution of the current sufficient to bring the operative
point back to P2.
If, on the other hand, a negative increase is applied to the current, the operative point D of the arc
corresponds to the operative point C of the machine, associated with greater voltage than
necessary, so that the current is increased and the operative point is brought back to P2 once
more.
Current
l Voltage
I
Uo
V
Icc Icc Icc Icc
79
If vice versa we analyse the functioning conditions of point P1, we find that the machine’s
characteristic, once the equilibrium is upset, moves the operating point ever farther away. P1 is
therefore an unstable operative point.
Fig. 135
From the above it can be seen that the arc’s operation is stable only when the external
characteristic of the machine is lower than that of the arc around the operative point.
V U0
Icc I
P2
P1
A
B
C
D
B
A
D
C
80
Welding processes using covered electrodes
The earliest electric arc welding applications date back to the end of the 19th century when the
covered electrode had not yet been invented.
The principle of electric arc welding was in fact discovered in 1892 whereas covered electrodes
appeared in 1908.
Bare wire welding has some considerable disadvantages among which:
- considerably difficult to start
- considerably difficult to keep going
- irregular weld metal deposition
- impossible to work in anything except flat positions
- loss of alloying elements that oxidise and volatise through the weld pool
- oxidization and nitrification of iron by the air present and therefore very poor mechanical
properties of the weld metal
- impossible to correct the weld pool analysis.
The list is long enough to exclude any interest in bare wire welding.
The coating on electrodes (Fig. 136) serves to eliminate these problems altogether or in part as well
as to supply the following functions: electrical functioning, shielding function, metallurgical function
and operative function.
Fig. 136
ELECTRICAL FUNCTION
Sodium and potassium silicates are always incorporated into electrode coatings because these are
highly emissive and facilitate ignition of the arc affording it good stability and making it possible to
use alternating current.
SHIELDING FUNCTION
The shielding function of the electrode coating is applied as a solid, a gas and a liquid.
Solid shielding is applied to the extremity of the electrode’s coating and as this melts after the core
does, it takes on a chalice shape and shields the material from contact with the air as it melts.
The elements contained in the electrode’s coating evaporate during welding creating a non-
oxidising atmosphere that is freed as much as possible of nitrogen. This atmosphere performs the
shielding function.
Liquid shielding is supplied by the slag that first wraps partially around the droplets of molten metal
and then, having a melting temperature lower than that of the core, coats the solidifying material
and shields it from contact with the surrounding air.
One element added to electrode coatings with the aim of reducing the fusion point is fluoride,
which however is de-ionising and therefore contributes to making the arc unstable.
Metal stick
Coating
Weld pool
Slag
Base metal Solid weld deposit
81
METALLURGICAL FUNCTION
The metallurgical function of the coating consists in the action of the liquid slag on the weld pool:
this can be summed up as: de-oxidising – purifying – inclusion of alloying elements.
De-oxidisation takes place through the ferroalloys (iron silicate and iron manganese) that reduce
the ferrous oxide that forms by bringing it to the surface in the form of silica and manganese oxides.
The purifying effect, typical of basic electrodes, is performed by calcium and manganese
carbonates that form composite elements with sulphur and phosphor; these composite elements
are brought to the surface of the weld pool as slag.
The inclusion of alloying elements usually takes place through the coating; it is in fact possible to
transfer elements from the coating to the slag and then to the weld pool with the aim of correcting,
where necessary, the analysis of the weld material deposited.
There are empirical tables from which the transfer percentages of various elements from the
electrode to the weld pool can be worked out.
These tables show, for example, that titanium cannot be transferred, so to weld a Ti stabilised type
321 steel requires using a niobium stabilised type 347 steel.
The alloying elements are included in the electrode covering in the form of powdered ferroalloys.
Electrode cores are always very pure and an analysis of the core will not give a valid indication of
the analysis of the material deposited during welding.
However, studies have been carried out recently on what are commonly called synthetic
electrodes, which supply the alloying elements mainly from their core and therefore have a neutral
coating; the advantages in producing electrodes such as these are considerable because the
same type of coating is used whatever the composition of the electrode.
Synthetic electrodes are generally those used in welding stainless steels.
OPERATIVE FUNCTION
The components of the coating have considerable influence on the characteristics of use of the
electrodes. Among these are working position (Fig. 137-138), electrical supply conditions,
penetration, shape and appearance of the bead, manageability of the electrode and the ease
with which the slag can be removed. This latter characteristic, for example, depends on the
content percentage of fluoride and titanium dioxide.
Fig. 137 Fig. 138
The higher the fluoride content, the more difficult it is to remove hardened slag: the edge of the
bead appears rough and uneven and the slag penetrates the irregularities more easily leading thus
to a mechanical anchorage that impedes its removal.
Titanium oxides make the slag dense and able to affect the conditions of solidification of the weld
pool through its surface tension leading to a bead with a good surface appearance. Titanium
oxides will also give the slag a shrinkage coefficient different to that of steel, thereby facilitating
separation of the bead during cooling.
Electrodes are marked with various symbols according to the classification standards involved.
The table given hereunder provides an example taken from UNI EN 499 standards that govern
covered electrodes for manual arc welding of non-alloy steel and fine grain steel.
82
Classification includes the properties ascertained using samples of solid weld metal and is based on
the use of 4 mm diameter electrodes.
Classification is split into eight parts indicated by symbols:
1. the first symbol identifies the type of product
2. the second identifies the class of resistance and elongation of the all-weld metal
3. the third identifies the toughness (impact properties) of the all-weld metal
4. the fourth symbol identifies the chemical composition of the all-weld metal
5. the fifth identifies the type of coating
6. symbol number six identifies the weld productivity rate and current type
7. the seventh symbol identifies the welding position
8. the eighth identifies the hydrogen content in the all-weld metal.
Classification is divided into two sections: mandatory symbols (1 to 5) and optional symbols (6 to 8).
We will now go on to see the eight elements making up the classification designation in detail.
1. Identification of the product: the letter E for covered electrode
2. Class of resistance and elongation
Symbol Minimum yield strength
N/mm2
Tensile strength
N/mm2
Minimum Elongation
%
35 355 440¸570 22
38 380 470¸600 20
42 420 500¸640 20
46 460 530¸680 20
50 500 560¸720 18
3. Impact
Symbol Test temperature for average minimum energy 47J
°C
Z Not required
A +20
0 0
2 -20
3 -30
4 -40
5 -50
6 -60
4. Chemical composition
Symbol Chemical composition %
Mn Mo Ni
No symbol 2,0 - -
Mo 1,4 0,3¸0,6 -
MnMo >1,4¸2,0 0,3¸0,6 -
1Ni 1,4 - 0,6¸1,2
2Ni 1,4 - 1,8¸2,6
3Ni 1,4 - >2,6¸3,8
Mn1Ni >1,4¸2,0 - 0,6¸1,2
1NiMo 1,4 0,3¸0,6 0,6¸1,2
Z Different composition according to application
Note: unless otherwise stated, values should be considered maximum.
83
5. Type of coating
Symbol Type of coating
A Acid
C Cellulosic
R Rutile
RR Thick Rutile coating
RC Rutile-cellulose
RA Rutile-acid
RB Rutile-basic
B Basic
6. Efficiency and current type
Symbol Yield % Current type
1 <105 CA + CC
2 <105 CC
3 >105<125 CA+CC
4 >105<125 CC
5 >125<160 CA+CC
6 >125<160 CC
7 >160 CA+CC
8 >160 CC
7. Welding positions
Symbol Welding positions
1 All positions
2 All positions except vertical in downward direction
3 Flat butt weld and for fillet weld flat and horizontal
4 Flat butt and fillet weld
5 All positions as under 3 + vertical down
8. Hydrogen content
Symbol Maximum hydrogen content ml/100 gr. of deposit
H5 5
H10 10
H15 15
EXAMPLE OF CLASSIFICATION: EN 499 E 42 4 B 4 2 H5.
We will now take a look at the American Welding Society (AWS)’s classification.
AWS 5.1 deals with the classification of covered non-alloy steel electrodes; the coded symbols
include the letter E followed generally by four figures (additional optional symbols refer to resilience
characteristics and hydrogen content).
The first two figures refer to the solid weld metal’s resistance to traction, expressed in kPsi (Kilo
pounds per square inch) where 1 kPsi = 7 N/mm2 . e.g.: E60XX - E70XX.
The second pair of digits conventionally indicates the type of coating/type of current; the main
ones are:
10 = cellulose ccpi
13 = rutile for dc and ac
15 = basic for ccpi
16 = basic for ac
18 = high yield basic
The RINA approval classes are given in Part D of the Rules.
84
Types of coating
Electrode coatings consist of a mixture of finely ground powders (oxides, minerals, carbonates,
ferroalloys, cellulose etc.) contained in a paste that is applied by extrusion to the electrodes’ cores.
With reference to UNI EN 499 standards, the principal types of coating are the following.
ACID COATING
These electrodes are easily handled and work well in any position, using both direct polarity DC
and AC; the weld pool is very hot and penetration significant.
Because of the presence of ferroalloys, the weld pool is also well deoxidised and has good
mechanical characteristics, although these deteriorate considerably at low temperature.
The hot weld pool and good penetration make the acid electrode particularly sensitive to the
formation of hot cracks when welding steels with high C (>0,25%) and impurity (>0,05%) content.
The coating is not particularly affected by damp, although the hydrogen content may sometimes
be rather high with the risk of micro-cracks forming when welding pieces of medium – thick
sections.
The use of these electrodes is therefore not advisable on thicknesses of more than 20-25 mm
without preheating.
This sort of electrode is widely used in constructions of light loaded steel structures.
CELLULOSE COATING
The fundamental characteristic of cellulose coating is its high content of organic substances
(cellulose).
The coating also contains limited amounts of deoxidisers and titanium oxide.
The purpose of this coating is to provide the weld pool with a mainly gaseous shield where there is a
reduced quantity of slag.
The shield is very efficient so the weld pool is well deoxidised and has good mechanical
characteristics.
Cellulose coated electrodes have a hot and fluid weld pool and high penetration; limited
production of slag makes them particularly suitable for welding in the vertical position.
Due to their high penetration these electrodes are in general used for the butt welds of pipes,
especially in the first pass when there is no support at the root; they are suitable both for vertical
down and vertical up welding techniques.
The hot weld pool and high penetration may cause the formation of hot cracks when the base
material has a high rate of impurity.
The weld metal deposited with these electrodes contains a very high percentage of hydrogen so
adequate preheating is necessary when welding thicker pieces of about 15-20 mm to avoid the
formation of cold cracks.
RUTILE COATING
It is similar to acid coating and differs only due to the considerable amount of titanium oxide
minerals. The weld pool is well deoxidised and the mechanical characteristics are only slightly
lower than those of acid electrodes.
These electrodes too have a hot and fluid weld pool and the slag, dense and viscous, is easily
removed.
They can be used with direct or alternating current and ease of ignition and stability of the arc are
their main characteristics; it follows that they are highly manageable and are suited to welding in
all positions.
Tendency towards hot cracking is more severe than with acid electrodes so particular attention
needs to be paid to the composition of the base material.
The higher hydrogen content can also lead to problems of micro-cracks in the welded zone.
These electrodes are therefore generally employed on thin sheets, finishing layers and in all cases
where the appearance of the final passes is of particular importance.
RUTILE COATING – THICK COATING
As those above but with a covering to core ratio of ≥1,6; the appearance of finishing passes and
ignition of the arc are improved.
85
RUTILE CELLULOSIC COATING
This consists of titanium oxide minerals and organic material it has characteristics between those of
rutile coating and cellulose coatings and is easier to manage than the latter.
It is generally used for welding pipes and thin sheeting in situ.
RUTILE ACID COATING
The operative characteristics of these electrodes are similar to those of acid covered electrodes.
A large proportion of the ferrous oxide content has, however, been replaced by rutile. These
electrodes, then, generally with a thicker covering, are suitable for welding in all positions except
the vertical down position.
RUTILE BASIC COATING
It consists mainly of calcium carbonate and oxide minerals with varying amounts of rutile.
It has characteristics similar to those of the basic coating.
BASIC COATING
The fundamental characteristic of the basic covering is a content of considerable quantity of
calcium carbonate together sometimes with magnesium carbonate and fluoride.
Moderate quantities of deoxidisers are also present and sometimes limited amounts of ferrous oxide
and titanium dioxide as well.
Fluoride is added to improve the melting process.
The weld pool is well deoxidised, has a low rate of impurity and very good mechanical
characteristics, often even at low temperature. As stated, the calcium carbonate and magnesium
carbonate form composite elements with sulphur and phosphor, which are brought to the surface
of the weld pool as slag.
Basic electrodes have a cool and therefore not-so fluid and moderate penetration weld pool; the
slag is hard to remove and the electrodes present certain difficulties in their handling which can
easily lead to porosity and are therefore not advised for thinner sections; they are generally used in
large scale boiler works and heavy duty steel structures.
During the manufacturing process these electrodes are dried at high temperature (around 450°C)
so their hydrogen content is very low, making them suitable for welding tempered steels with a
reduced risk of cold cracks forming.
The covering is, however, very hygroscopic so these electrodes must be stored in airtight
packaging and partial pre-drying must be carried out at 400°C for an adequate length of time
after which they need to be kept at about 100°C until they are used.
Edge preparation
Before welding it is necessary to prepare the edges to be joined in order to ensure good
penetration and ease of welding needed to obtain a healthy joint. Preparation work varies
according to the type of connection, the thickness of the pieces, whether or not both sides can be
reached, the welding position, and the project requirements.
In general for butt-joints reference may be made to the UNI EN ISO 9692-1 standards.
Butt-joints
a. Straight edge preparation: this is used on thickness < 5 mm; when a backing on the root side is
not used, adequate gouging must always precede the second pass (Fig. 139).
Fig. 139
1÷3 ≥4
86
b. “V” preparation: can always be used for thickness > 2-3 mm; for thickness > 15 mm where both
sides can be reached, the “X” preparation might be preferable. Where possible it is preferable
in any event to do the second pass on the root side, after adequate gouging. A backing on the
root side (d> 4 mm) can also be used in this type of preparation.
Fig. 140
c. “X” preparation: may be symmetrical or asymmetrical. Root gouging must always be done (Fig.
141).
Fig. 141
Other types of preparation work can be performed (Fig. 142); among these, mention should be
made of the ½ V (horizontal welding position) and the “U” preparation (very thick sections).
Fig. 142
"T" joints
They can be made with partial and full penetration (Fig. 143).
60÷70°
2÷4 0÷2
2÷3
3
60°
1/2
1/2
60°
2÷3
3
60°
1/3
2/3
60°
87
Fig. 143
Welding defects related to the weld execution
Defects can be external or internal.
Among the external defects that can be detected during visual inspection or non-destructive
surface tests (LP-MT) we have:
Undercuts that appear as indentations on
one or both sides of the weld bead. Too high
a current or too long an arc (poor welder skill)
usually leads to these defects (Fig. 144).
Fig. 144
Incorrect profile, which is generally caused by an irregular or unsuitable welding speed, a wrong
movement of the electrode or the use of incorrect current (Fig. 145).
In butt-joints that do not have a second pass on the root, the defect might also be present on the
root.
88
Fig. 145
Crater defects can take place each time a pass is finished, especially if thick electrodes with a high
penetration characteristic are used (Fig. 146). These can also cause cracking. The remedy is to
distance the electrode from the piece only after dwelling on the last weld pool a little longer.
Fig. 146
Among the internal defects that can be detected with non-destructive tests such as RX, UT and in
some cases MT we have:
Incomplete penetration, which is
characterised by a lack of fusion on one of
the edges (Fig. 147) or at the root of the bevel
(Fig. 148) or at the corner of a fillet joint (Fig.
149).
This is a very serious defect because not only
does it reduce the resistant section of the joint
but it also allows a notch to form in the same
place, which can lead to cracks and fracture
and in some cases to corrosion. The causes
of this defect are many: incorrect welding
parameters, excessive diameter of the
electrode, insufficient gouging on the root
side and a lack of welder’s skill.
Fig. 148
Fig. 147
Fig. 149
89
Internal cavities: generally irregularly shaped cavities that may also contain slag in which case they
are identified as “slag inclusions” (Fig. 150 & 151). The cause of these can be: an incorrect pass
sequence, in which case the cavities are usually hairline, or an irregular profile left by previous
passes, or a wrong movement of the electrode, in which cases the cavities can have a variety of
shapes.
Fig. 150 Fig. 151
Lack of fusion between the base metal and the weld metal (Fig. 152 & 153). Although in electric
arc welding the fusion always takes place in the electrodes and in the base metal, lack of fusion
can happen when the arc is too long or the intensity is too low, such that the first droplets of weld
metal, at the start of the pass, reach the base metal (which has a greater thermal capacity) not
yet molten but extremely hot and therefore oxidised; or when the arc deviates in the bevelled area
due to the magnetic blow effect or the irregular fusion of the coating.
This defect is hard to detect with X-ray tests.
Fig. 152 Fig. 153
Pores are the inclusion of gas still under pressure within the material; they can have a spheroid or
elongated shape and may be joined together in clusters or be isolated (Fig. 154 & 155); sometimes
they appear at the surface of the bead, and when they are very small they are identified as
general porosity.
90
Fig. 154 Fig. 155
Cracks are a very serious defect but depend on the metallurgical characteristics of the base and
weld materials rather than on the technique employed (Fig. 156 to 159). It is useful to bear in mind
in any case that the welding technique used can favour the formation of cracks. We have already
spoken about cracks that propagate from a crater. Using high currents or even too thin a first pass,
incisions and indentations can lead to the formation of cracks.
The presence of other defects such as blowing, slag inclusion or a lack of penetration can also
contribute to cracking.
Fig. 156 Fig. 157
Fig. 158 Fig. 159
91
Submerged arc welding
This process originated in America and offers a much higher working speed than manual arc
welding (Fig. 160).
Fig. 160
A coil of continuous bare wire substitutes the short covered electrode so that down time
connected with changing electrodes is eliminated.
The arc spans between the wire and the base material
and is shielded by a layer of powdered granulated flux
which is gravity fed from a hopper to the weld area; as the
arc is completely covered by the layer of flux, it is hidden
from view (submerged arc) (Fig. 161) and the operator
does not need to shield his eyes.
The submerged arc can be used in flat welding, flat-
horizontal welding and with appropriate caution in
horizontal welding; the process can be automatic or semi-
automatic (the latter has been almost entirely abandoned
nowadays).
Fig. 161
The high current intensity allows a
considerable amount of weld metal and
parent metal to be melted with high dilution
ratio and the possibility of welding very thick
sections (Fig. 162).
Fig. 162
92
The welding device is composed of: a wire feed mechanism; a rocket on which the wire is coiled;
the current brushes in proximity to the arc with the purpose of reducing voltage droplets and the
Joule effect; the current, voltage and pass speed control and monitoring commands; a hopper
containing the flux and safety devices (Fig. 163 & 164).
Fig. 163 Fig. 164
The welding speed can be obtained either by the movement of welding head or in some cases of
the piece being welded and it has to be adjusted each time depending on thickness and other
welding parameters used.
The current generator can be dc (dynamo and transformer plus rectifier) or ac (transformers).
Using dc the arc is more stable and the load is evenly distributed on the line.
Using ac we have lower transformer cost, lower no-load consumption, less maintenance and a
weaker influence on magnetic blowing.
Direct current is normally used for intensities of less that 700 A whilst alternating current is used for
intensities in excess of 1200 A; in the interval between 700 and 1200 A it makes no difference
whether ac or dc is used.
The factors that mostly influence the shape and quality of the weld bead are, in approximate order
of importance: welding current – welding voltage – welding speed.
Of lesser importance: diameter of the electrode, length of stick-out, type of flux and wire.
- an increase in the current leads to higher penetration and higher weld reinforcement while the
width of the bead remains more or less constant
- an increase in the voltage leads to the bead widening while penetration stays remarkably
constant. As the voltage increases, so does the consumption of flux.
- an increase in speed leads to a diminution of the weld reinforcement and of the width: tripling
the speed, weld reinforcement and width are halved.
- an increase in the diameter of the wire reduces penetration. That is, with the same current,
voltage and speed, a smaller wire will give greater penetration. The width of the bead is not
influenced by the diameter of the wire.
- stick-out length, defined as being the distance between the electrical adduction contact of
the welding current and the tip of the wire being fused, takes on particular importance with
moderately high current densities (≅ 102 A/mm2); through the Joule effect the deposition rate
increases (even by as much as 50%) but penetration decreases. The effect on the weld bead of
increasing stick-out is the same as reducing the voltage. This parameter is to be controlled
carefully because it can lead to morphological characteristics differing from those expected
and desired.
- the type of flux and wire can also influence the operative conditions: for example under
identical conditions a stainless steel wire, having a greater resistance, is fused quicker.
Wires and flux have various chemical compositions just as covered electrodes have and in the
same way the flux has identical functions in the weld pool and the fusion area. Therefore, the first
93
flux began to be made by grinding the coverings of electrodes. This, however, posed one problem
in that the gas released from the flux during fusion often got trapped in the powder, which was
impermeable to it, so that porosity was formed in great proportions.
Weld flux (Fig. 165) requires specific manufacturing processes; it is produced in two main types:
Pre-fused or neutral flux: the components making up
the flux are mixed and then fused and ground. Fusion
leads to reactions between the various components
and in particular to the liberation of gas.
Clearly, by suppressing all reactions during the welding
process, the flux behaves as a neutral flux providing a
shield to the weld pool mainly through mechanical
cover.
This flux has no chemical-metallurgical function so
there is a need for additional “active wires”, i.e. wires
that contain alloyed deoxidising elements, such as
manganese usually. Pre-fused flux is not particularly
hygroscopic, that is it does not absorb humidity.
Fig. 165
Agglomerated or active flux: obtained by mixing the required components in a finely ground
powder which is then agglomerated by oven treatment using appropriate binders.
The agglomerate is then ground down and sieved according to a granulometer rough enough to
ensure gas permeability. Compared to pre-fused flux, agglomerated flux keeps all its reactions for
the welding process and is therefore referred to as being “active”. This flux has a chemical function
that is an active participation in the analysis of the weld pool. Therefore, pure wires rather than
alloy wires can be used, leaving it to the flux to work on the composition of the weld metal
deposited. The elements to be found most frequently in flux are: SiO2 - Al2O3 - CaO - MgO - CaF2
and, in the case of agglomerated flux, ferroalloys as well.
The proportions between the different ingredients vary according to the type of flux.
As regards wires (Fig. 166), there are commercially available C
and C-Mn wires for welding steel that are characterised by their
manganese content. With regards to the flux, the wire must be
chosen in relation to the activity of the flux. The choice involves
the most appropriate combination of flux and wire. The alloying
elements whose effects are to be monitored with the greatest
attention are Si and Mn.
During experiments it was found that the analysis of the weld
pool with a ratio of Mn/Si greater than 2,5 is necessary. This
reduces the risks of cracking.
As for covered electrodes, there are various classification
standards for wire and flux combinations. Given the complexity
of the classification standards we will limit ourselves to
considering the classes approved by RINA, for which reference
should be made to Part D of the Rules. Fig. 166
The characteristics of submerged arc welding that are of particular importance, especially when
high current intensity is used, are:
- the high percentage of base metal in the deposited weld metal (dilution ration); it is important
that the more harmful impurities contained in the base metal (sulphur, phosphor) are low in
content
- the considerable quantity of slag and the high risk of exchange with the weld pool means that
the flux humidity is to be considered in particular; the lower the temperature at which it was
manufactured the poorer the state of preservation will be and the greater the steel’s
susceptibility to cold cracking
94
- the large amount of heat input influences the crystalline structure in the fuse area and the heat
affected zone. In particular, welding using high current and low welding speed leads to a
dangerous drop of the mechanical properties, especially of the impact properties, in the heat
affected zone area and the fusion area.
The following short list details some aspects and techniques of submerged arc welding.
Techniques
Submerged arc welding was invented as a high-speed process capable of producing
considerable quantities of deposited metal per pass on moderate and thick sections; as already
seen, this can happen to the detriment of the final mechanical properties of the welded joint.
Generally speaking it can be said that a technique called “high penetration” (Fig. 167 & 168) is
used for steelwork that does not require high metallurgical quality welding. This exploits the
productive capacity of the process but the metallurgical results are in general moderate. In boiler
work on the other hand, where considerable metallurgical quality is needed, welding is done using
multiple passes (Fig. 169).
The evolution of base materials and weld metals contributes to an ever-higher exploitation rate of
the process’s high productivity characteristic.
Fig. 167
Fig. 168: single pass on 10 mm plate
Fig. 169
Fig. 170 shows the “tandem” technique: the two weld heads ensure an even greater deposition
rate.
Fig. 171 shows the technique with the strip electrode, used for electroplating.
Fig. 172 & 173 show a type of panel line widely used in shipyards to weld deck and bulkhead
plating.
95
Fig. 170 Fig. 171
Fig. 172 Fig. 173
Preparation of joints and methods of execution
Butt-joints
a. Preparation of straight edges. This is employed particularly with the high penetration technique
to obtain joints with full penetration after just two passes on the opposite sides of the joint.
Contrary to the technique employed in manual metal arc welding, gouging is rarely done
before the second pass in this technique. The thickness that can be welded using this
preparation job can, in certain exceptional cases, exceed 15 mm.
b. Preparation with bevel. The preparation can be “X” or “Y” (dimensions of the face 6-7 mm),
which for thicknesses up to 16-18 mm can be done using the two-run technique (single pass per
side); for greater thicknesses the bevel is usually filled with multiple passes or a better preparation
can be the “X” bevel (symmetrical or asymmetrical). With these latter preparations the multiple
pass technique is generally used.
Other types of bevel can be used in the one side welding, using a support that might be a series of
base runs done using a different welding process, support plates made of strips of steel compatible
with the metal to be welded, copper supports, which must not be brought into the fusion, with
96
appropriate side channels, and flux beds that are adequately pressed against the reverse side of
the joint by pneumatic devices.
T-joints
The same considerations given for the manual process apply; devices using two weld heads to
weld both sides of the vertical beam at the same time may be employed.
Submerged arc welding can sometimes employ devices with two or more wires having single or
multiple head with the evident purpose of increasing productivity.
Welding defects related to the execution
In general the same considerations given for the manual process apply; particular attention should
be paid to the possibility of lack of penetration in the two-run technique, a defect that is not easy
to detect from an X-ray test; the cause of the defect can be found in the inadequate welding
parameters, the diameter of the wire used, edge preparation and, especially where straight edges
are used, an incorrect alignment between the two passes.
97
MIG and MAG welding processes
In this type of process the electrode is made
up of the weld wire which, being fusible, has
to be fed with continuity.
The welding devices are therefore analogous
to those we have already looked at in the
submerged arc welding process, except that
a shielding gas replaces the flux (Fig. 174-175-
176-177-180).
Fig. 174
Fig. 175
Fig. 176
98
The speed with which the wire is fed directly
controls the welding current.
The electrode welding process, when it was
first used, employed almost exclusively argon
as a shielding gas and its practical
application was limited to welding stainless
steels, nickel, copper, aluminium and
aluminium alloys. The applications for carbon
steel welding would be too expensive given
the high cost of the gas.
Fig. 177: tandem torch
The welding process that uses consumable electrodes and argon shielding (or other inert gas) is
conventionally referred to as MIG (Metal-arc Inert Gas).
In relatively recent times a process has been invented in which the shielding gas is carbon dioxide;
the cost of this gas, considerably lower than that of inert gases, has allowed the process to be
extended also to welding carbon steels; carbon dioxide is, however, chemically active at the
temperatures met in correspondence with the arc so it was necessary to design wires with
particular chemical compositions suited to performing an adequate deoxidising role.
Conventionally, the process using CO2 as a shielding gas or other gases that are active (UNI EN 429
standards define as active any mixture that contains even a minimal percentage of active gas,
usually CO2 or O2; see Fig. 178) is referred to as MAG (Metal-arc Active Gas).
Compared to the manual welding with covered electrodes, the continuous wire process presents
the following advantages:
- continuous feed of the weld metal and therefore increased productivity
- practical absence of slag (when solid wires are used)
- visibility of the weld pool (when slag is absent) with a consequential better control of the weld
- a higher current density that allows high deposition rates
- lower risks, when using solid wires, of cold cracking (hydrogen cracking) in materials that are
susceptible to this; the humidity content of the shielding gas can be very low
Again compared to welding with covered electrodes, the continuous wire process can, however,
present some limitations or at least disadvantages:
- the device is more complex and therefore more costly and difficult to move in tight spaces
- the weld gun is bulky and makes the process less suitable for joints in spaces that cannot easily
be reached
- when solid wires are being used there is no possibility of purifying the weld pool, something
typical of basic electrodes, so it is possible that hot cracks will form even if steels with only
moderately high rates of impurities are being welded
- a certain amount of caution is required to avoid draughts of air reaching the work area so as
not to compromise the gas shielding
Compared to the submerged arc process, of which this is a natural development, it offers a
considerably lower productive rate but allows direct control over the weld pool and has the main
advantage of being suitable for use in all positions.
Main shielding gases
Argon is a single atom gas that cannot therefore be split or react with any other element; it has a
high atomic weight, which makes it particularly active in removing rust film when welding light
alloys and manganese. As it is almost insoluble in the weld pool, the risks of this gas ending up in
the bead are very low. As mentioned earlier, it is used for welding aluminium alloys, nickel, copper
and stainless steels; for the latter it can be used in a mixture with small percentages (generally ≤
2,5%) of CO2 or O2.
99
Carbon dioxide is an inert gas at ambient
temperature but at the temperature of the
arc it splits into carbon monoxide and
oxygen; for this reason it has an oxidising
effect on the weld pool. Carbon dioxide also
has a carburising effect on the weld pool:
CO2 + C → 2CO. If the C content of the weld
pool is low and the CO provided by the
chemical reaction in the weld pool is high,
the reaction takes place from right to left, so
there is an enrichment of carbon in the weld
pool: this practically always happens when
the C content in the weld pool is lower than
0,07%; in any case (even when the C content
is considerably higher), the C content of the
weld pool shielded with CO2 is set as a
percentage of between 0,07 and 1.
Fig. 178
The use of CO2 is consequently not advisable when welding materials where critical alloying
elements (C-AI-Ti-Si-V-Mn-Nb) have to be controlled; carbon dioxide is therefore not suitable for
welding high alloy steels or steels with a controlled C content (e.g. quenched and tempered) to
avoid risks of oxidising and carburising the weld pool. Other problems that might arise when using
CO2, in particular if the welding equipment is not properly controlled:
- formation of porosity in the weld metal (FeO + C→ Fe + CO); the carbon oxide that forms from
this is trapped in the weld pool, leading to internal porosity. To avoid this reaction it is necessary
to add some deoxidising and killing elements such as Si and Mn to the weld pool.
- metallic splatter along the edge of the bead. This is caused by the way the weld metal is
transferred in the weld pool; this problem can be avoided by using high current densities –
however, because of the large weld pool the process cannot be applied to welding thin plates.
Adopting particular electrical parameter values the way the droplets are transferred can be
modified while keeping the current density at a value allowing thin plates to be welded.
- limitation on position welding. This is a result of what was mentioned above: the large weld pool
makes it possible to use it in flat welding only. This problem has also been resolved by modifying
the way the metal is transferred along the length of the arc.
To get over these difficulties various gas mixtures have been proposed: those used normally in the
field of welding C and C-Mn steels are those composed of Ar + CO2. Mixtures with percentages of
CO2 variable from between 15 and 25 are generally used.
Fig. 179 illustrates the effect of different shielding gases/gas mixtures on the shape of the weld
bead.
100
Fig. 179
Wires
The importance of the wire in gas shielded welding is quite remarkable: the chemical composition,
the manufacturing method, the presence of flux (flux-cored wires), the surface state and the
spooling are very important factors that affect the welding.
Surface state: must be excellent, with no
traces of grease or humidity in order to avoid
the possibility of porosity, wormholes, cracks,
reduced electrical contact.
Spooling: because the speed with which the
wire is fed regulates the intensity of the
current, if spooling is not done properly it can
lead to serious fluctuations of the current with
consequential risks of irregularities in the
bead, sometimes unacceptable.
Fig. 180
SOLID WIRES
Solid wires for MAG welding (C and low alloy steels), i.e. with a gas or mixture with a more or less
oxidising effect, are characterised by a fairly high Si content which is energetic and deoxidising. Si
is generally present in percentages of from 0,3% up to 1,2%. Higher values could lead to a brittle
structure and increase the tendency to hot cracking.
Mn works as a deoxidiser, as a desulphuriser and as an alloying element. In these applications it is
usually present in values of between 0,8% and 2%.
Careful preliminary tests have shown that the loss resulting from the deoxidising effect is on average
50% for Mn and 60% for Si.
Solid wires for welding medium alloy steels may contain elements that are integrally transferred to
the weld metal if the shielding gas is inert or partially active.
Solid wires for welding alloy steels (stainless steels) generally have the same chemical composition
as the base material, possibly with a low C content.
The most commonly used diameters of solid wires are 0,8, 1, 1,2 and 1,6 mm.
Fig. 181 gives shows the chemical compositions of solid wires for welding carbon and carbon-
manganese steels provided for in the RINA Rules.
101
Fig. 181
FLUX-CORED ELECTRODES
These have a hollow centre that is filled with flux that can have the same purpose as the covering
of electrodes (Fig. 182). Certain wires of this type can be used without shielding gas.
The external metallic surface is almost always
made of soft iron and the active and alloying
elements are contained within the flux inside.
The characteristics of the flux may be basic,
rutile, rutile-basic and less often acid or
neutral. There are various types of wire
produced by bending metallic strips or
drawing small metallic tubes.
Compared to solid wires, flux-cored wire
allows a faster rate of deposition, higher
penetration and the possibility of welding
materials with traces of oxide on their surface
or alloy steels (sometimes even stainless steels)
using CO2 protection.
Fig. 182
On the down side there is a greater consumption of gas, the problem of the slag that has to be
removed and the generally higher cost.
Transfer of the weld metal during welding
The way the droplets of metal are transferred along the length of the arc depends on the diameter
of the wire, the type of generator and shielding gas used as well as the electric welding parameters
(A, V). On the basis of these parameters transfer may take place by:
- spray-arc
- dip-transfer or short-arc
- pulsed-arc
Experimental tests have shown that the minimum deposition rate must be about 50
droplets/second to obtain a bead with satisfactory characteristics.
With low intensity current, the droplets grow at the tip of the wire until they reach a volume such
that they fall off the wire under their own weight with a minimal frequency. On reaching critical
current intensity levels, which depend on the diameter of the wire and the type of shielding gas,
the dripping frequency increases and transfer takes place not through gravity but by virtue of
particular electrodynamic forces (magnetic contraction effect, plasma flow etc). In any case, only
by using inert shielding gases will the flow of free droplets be axial and therefore properly
governable (with positive effects on the quality of the bead and with little weld splatter); with
active gases or mixtures, while drip frequencies higher than 50 droplets/minute can be reached,
the free flow transfer is never regular.
102
In any case, in this scenario – with a drip
frequency greater than 50 d/m – the transfer
is said to be “spray” or high intensity type (Fig.
183). With spray-arc transfer we have a
quieter arc, penetration increases, the bead
appearance is better and spatters almost
disappear. As we have seen, the electric
welding parameters that influence the type
of transfer depend on various factors, but as
a guide we can say that with C steel and a
1,2 mm diameter solid wire with CO2 shielding
gas, transfer will be by spray-arc where the
current is above 200 A and the voltage
above 22V.
Fig. 183
If, starting from the spray-arc function, we
lower the current to what is referred to as the
transition current level, the function becomes
globular and the working conditions become
prohibitive. If we reduce the voltage too,
however, the arc becomes shorter and the
droplets forming create a continuous liquid
bridge between the tip of the wire being
fused and the weld pool. The current tends
rapidly towards the short-circuit value and
cuts the droplets by the neck through the
neck down effect. The arc is re-started,
another drop forms, a new short-circuit is
established and so on (Fig. 184).
Fig. 184
In order for the short-arc to work properly, without sticking or excessive spattering, the following
conditions must be met:
- current intensity between 70 and 180 A
- voltage between 16 and 25 V
according to the diameter of the wire, the type of generator and shielding gas.
A correct short-arc regime requires not less than 100 short-circuits a second.
A practical comparison between spray-arc and short-arc (Fig. 185) leads us to the following
considerations: the spray-arc regime can always support very powerful arcs with high current
density suitable for flat welding moderate and thick sections. For lesser thicknesses, the sheet would
burn through and, working on moderate thicknesses in the vertical position, the weight of the weld
pool would bring it down. The short-arc regime is cooler because the current and voltage are
considerably lower. This makes it possible to weld in the vertical position and to work on thin plates.
The low amount of heat supplied in the short-arc regime leads to two serious worries: sticking
through the non-fusion of the base material and the formation of blowholes through an overly fast
solidification of the weld pool.
In between the excessive power of the spray-arc regime and the often insufficient power of the
short-arc regime, there is a gap that precludes MIG – MAG processes from many applications.
40
30
20
10
400 300 200 100
short
globular
spray
V
I
103
Fig. 185
A third transfer method, called pulsed-arc, has been put forward as a way of filling this gap. This
consists in feeding the wire with a current that is lower than the transition current and applying a
current pulse on top of it in order to rise periodically above the transition threshold (Fig. 186). Two
generators running in parallel supply the arc: one dc generator maintains the background current
and a pulse generator creates the current peaks. The intensity of the pulses, their duration and
frequency, their shape and the level of the background current are variables that influence the
length of the arc, the speed of fusion, the section of the bead and the spatter phenomenon. The
possibility of working with a current that is on average below the transition current level whilst still
maintaining a spray-arc transfer means it is possible not only to weld thin plates but also to weld
greater thicknesses in positions different to the flat one. With respect to the traditional spray arc,
pulsed-arc transfer allows aluminium to be welded reducing the rate of wire fusion by about a half.
Fig. 186
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Preparation of joints and methods of execution
In general the same preparations are used as for the process with covered electrodes; sometimes,
given the good workability of the electrode wire in the bevel itself, the caulking angle may be
reduced to 45°-50°. In welding with a flux-cored wire, substantial use has recently been made of
ceramic supports enabling welding from one side only with evident economic advantages (Fig.
187 & 188); in this case, in any event, special attention needs to be paid to the preparation of the
joint, so that the distance to the root is from 4 to 6 mm.
With MIG and MAG processes, welding in all positions involves relatively easily methods of
execution; however, using fully basic flux-cored wires, generally welding in other than flat and
sometimes horizontal position is not possible.
104
Fig. 187 Fig. 188
Welding defects related to the weld execution
The defects are largely those already described for the process with covered electrodes.
Blowholes are generally caused not only by an arc of excessive length, but also by inadequate
protection of the weld puddle; accordingly, special attention must be paid to the flow of gas
(generally 10 times the diameter of the wire, in litres per minute, but in the presence of particularly
long sheaths account needs to be taken of a possible drop in flow). It is also necessary to exercise
care so as to prevent air currents from hitting the zone of the arc and impairing the effectiveness of
the gas protection.
We have already mentioned the possibility of lack of fusion of the base metal (at the edges of the
bevel) or also of the weld metal already deposited (lack of fusion between passes) caused by use
of particularly low currents obviously combined with the welder’s lack of skill. This defect is very
dangerous and hard to identify using X-ray tests but it can be detected by means of UT tests. In the
event of use of flux-cored wires, lack of fusion is less common insofar as the operation is carried out
with lower welding parameters than solid wires; in any case this defect is often accompanied by
slag inclusions, which make it easer to detect using X-ray tests.
105
TIG welding process
The abbreviation TIG stands for Tungsten Inert Gas welding, or the use of tungsten electrodes in an
inert gas atmosphere.
The arc is started between a refractory tungsten electrode and the edges to be welded while a
flow of inert gas (usually argon) comes from the electrode torch enveloping the tip of the
electrode, the arc column, the tip of the welding wire and the weld pool (Fig. 189 & 190).
The tool referred to as the electrode torch is a refractory tipped pipe at the centre of which the
expendable tungsten electrode is inserted, and a grip that is connected to the flexible supply
sheath (Fig. 193 & 194).
The flexible sheath carries current, the shielding gas and, where used, cooling water to the torch.
The inert gas is taken from a cylinder with a reduction valve and a flow meter. It is therefore possible
to keep the consumption of gas to within the limits needed to obtain a shield. As a guide we could
say on average 10 l/minute.
The technique of execution is relatively simple and is in appearance like the oxyacetylene method.
The torch serves solely to support the arc while the weld metal is fed in separately in the form of
stick wires. Consumption of the tungsten electrode is almost zero (Fig. 191 & 192).
Fig. 189 Fig. 190
Fig. 191 Fig. 192
106
Fig. 193 Fig. 194
Electrical supply
The choice of current to supply the arc is strictly related to the nature of the material to be welded
and more precisely we must distinguish between two groups of metals as follows:
A) B)
- Common and alloy steels - Aluminium and aluminium alloys
- Copper and copper alloys - Manganese and manganese alloys
- Nickel and nickel alloys - Aluminium-bronzes (Cu-Al alloys)
- Titanium
For the materials belonging to group A, direct current with direct polarity (negative pole to the
electrode) must be used.
With direct current in fact, the arc is more stable and, with direct polarity, the flow of electrons from
the electrode to the weld pool increases penetration and accelerates execution. With direct
current and reverse polarity the tip of the electrode would be bombarded with electrons leading
to extreme heating and fusion of the electrode despite it being refractory. Fusion of the electrode,
apart from the waste involved, would also lead to unwanted inclusion of tungsten in the weld pool.
The materials belonging to group B have in common that they are easy to oxidise so that their
surfaces are always coated with a thin layer of oxidisation. Before going on to weld, the bevel
surfaces must be chemically or mechanically deoxidised: despite this and while benefiting from the
gas shield, during welding oxides shows up on the surface of the weld pool. Both Al and Mg oxides
are refractory (they melt at above 2000°C) and their density is higher than that of the metal. From
a practical point of view the film of oxide that remains solid on the fused metal prevents the welder
from seeing the fusion and there is therefore a risk of burn through. In addition, the solid film wraps
around the droplets preventing their union and leaving inclusions. Removing these oxides is
possible electrically, using direct current with reverse polarity; the phenomenon is due to the flow
towards the weld pool (cathode) of positive ions of inert gas that have far greater mass than
electrons and therefore more kinetic energy. This action is called ionic sandblasting because it
reproduces, on a reduced scale, the operation of sandblasting performed to remove layers of
certain metal surfaces.
The problem would be solved were it not for the inconvenience of having the electrode
bombarded with electrons, which as we said earlier leads to the tungsten wearing down rapidly.
Using alternating current for each semi-period of the voltage wave, the electrode is positive and
that allows the film of oxide to be removed whilst in the other semi-period the electrode is negative
and that serves to limit heating of the tip. It is in this way possible to exploit the advantages while
reducing the disadvantages of the two types of ac electrical supply (Fig. 195).
107
Fig. 195
The weld metal use with TIG welding is very often similar, though not necessarily identical to, the
parent metal; the material of which the stick wire is made is in fact often provided with limited
quantities of deoxidisers. The choice of weld material, for every application, depends on the
metallurgical compatibility and takes into account of course the mechanical properties, resistance
to corrosion, electrical and thermal conductivity etc that are required from the welded joint.
Attention should be given to the fact that the stick wire, the welder’s gloves and of course the
bevel must be scrupulously clean before starting to weld.
Joint preparation and methods of execution
The TIG process can be used in every welding position; the arc is quiet and weld pool and fusion of
the edges are easy to control, the first passes have a particularly good appearance, and therefore
there is no need for a second run on the reverse side. The method is therefore used for first passes
on pipes, even of small diameter, and for welding particularly thin joints. Its low productivity and
high cost make it ill-advised in filling and finishing passes. It is practically impossible to substitute and
still obtain equal results in welding stainless steels, titanium, aluminium alloys, copper and nickel.
With regard to preparation of the edges, this has to be done with particular care and attention.
Due to the size of the torch, when performing first passes on rather thick sections, bevel angles must
be more open than those used in processes already discussed (≥ 70°).
In some cases where welding thin sections, an appropriate set-up of the edges (generally zero root
gap) allows the process to be used without adding weld metal, bringing the base metal itself to
fusion.
Welding defects that depend on the execution
Gas shielding and the use of solid electrodes can lead to inclusion of slag; when referring to
blowholes and porosity, the same considerations given for MIG/MAG welding are valid here as
well.
We have already mentioned the possibility of tungsten inclusions in the weld metal; because these
are sharp fragments of a hard metal inside a weld joint, tungsten inclusions are generally
considered serious defects. In an X-ray test they show up as very clear spots.
108
5. Weldability of materials
Weldability is any given material’s attitude in lending itself to a welded union that provides
predicted properties.
The definition of weldability established by the International Welders Institute is as follows:
“A metallic material is considered weldable at a certain degree by means of a given process and
a certain type of application when, after setting up the precautions necessary, it lends itself to the
realisation of constructions made up of elements for which it is possible to guarantee metallic
continuity through welded joints which, through their local properties and global consequences of
their presence, satisfy the quality criteria set out as the basis of the judgement”.
Certain elements of this definition need looking at. Weldability is not an element that can be
defined in an absolute way but only by degrees (degree of weldability) and these degrees are
valid for any type of material, according to the welding conditions set up. For example, each type
of steel will have its own degree of weldability in each process used, and this will impose
precautions that are all the more important the lower the degree of weldability is.
For the most part when one talks of weldability reference must also be made to the aims
established in the construction of the welded work, so weldability is a property conditioned by the
properties of the welded joint that is designated as the basis for the judgement, considering them
necessary to reaching the goal of the welded work performing correctly.
The definition also considers local properties and global consequences of the welded joints
because:
- each welded joint has a fused zone and a heat affected zone so the material must be weldable
in such a way as to create within it metallurgical structures providing the properties required of
them; in particular, there must be no serious defects like hot cracking, cold cracking or lamellar
tearing
- the welded joints must be such that the properties of the construction overall are not impaired,
i.e. to guarantee global safety of the construction; in this respect it is the properties of resilience
of steels that play a major role.
109
Mild steels Mild constructions steels are steels with a carbon content of not more than 0,23% and in which the
percentage of other alloying elements is very low and consequential mainly to manufacturing
needs.
FE 360 and Fe 340 type steels fall into this category (in shipbuilding, standard hull steel with Rm ≥ 400
N/mm2, normalised or fine grain normalised – Fig. 196).
The easiest to weld with all processes, mild steels do not pose excessive problems of hardening in
the heat affected zone and the metallurgical problems to keep an eye on are therefore hot
cracking, lamellar tearing and ruptures of brittle metal.
Lamellar tearing can happen when thick sections are being welded. The phenomenon is normally
tackled through precautions during the design and execution stages.
Hot cracking is to be feared in particular when there may be a high level of impurities; also, in the
case of effervescent steel, there may be considerable concentrations of impurities in the areas of
segregation: the Mn/Si and Mn/S ratios present in the weld puddle are indicative of a greater or
lesser susceptibility to hot cracking. The deoxidising and desulphurising properties of manganese
are in fact known, as is the brittleness caused by a high level of silica. When these values reach
differences of Mn/Si > 3 and Mn/S > 20, safe conditions can be considered as being achieved.
In manual stick welding, the use of good basic electrodes is generally precaution enough to avoid
the formation of hot cracking when high levels of impurity are to be feared. Therefore, considering
the negligible tendency to hardening in the heat affected zone, cold cracking is generally not
something to worry about.
Where the phenomenon of brittle ruptures is the problem (low operating temperatures, large
constructions, considerably thick sections), the steels to be chosen must guarantee an adequate
level of resilience (> 27J) at the minimum operative temperature.
All the welding processes can be employed without particular precautions; moderate pre-heating
might be suggested where there are very thick sections to weld or where the ambient temperature
is very low.
Heat treatment is only employed for even higher thicknesses and in relation to the type of steel, the
need for it being more evident the higher the steel’s resistance to traction.
Fig. 196
110
Low manganese alloy steels
This designation is given to low alloys steels which, as well as carbon, contain 1 to 1,5% addition of
manganese with the purpose of improving the resistance; breaking loads of around 500 N/mm2 are
thus achieved (high resistance steels with Mn 0,90÷1,60% - Fig. 197).
Fig. 197
Manganese, like carbon, favours hardening of the steel; however, its action in this direction is very
bland. Additionally, these steels have a carbon content in the order of 0,2% so their hardening
ability is, all told, only slightly greater than that of mild steels. High resistance steels are followed
closer during manufacturing so they are always normalised or, at worst, semi-normalised.
When one wishes to increase the resistance further still, without making welding too difficult, it is
possible to add small quantities of alloying elements such as vanadium, titanium and chrome,
making micro-alloyed steels. The latter have little hardening ability; however, welding of the most
resistant steels should be done with a certain amount of caution.
Although the content in carbon and impurities is generally limited to sufficiently low values, the
danger of hot cracking might still exist. It is therefore generally wise to use processes and weld
metals that are able to purify the weld pool, especially in the case of welding of large thicknesses.
With regard to hardness and brittleness of the heat affected zone, and therefore the danger of
cold cracking, the most significant indications can be taken from the TTT curves of the CCT
diagram where available. These might also show whether or not it would be necessary to do some
pre-heating, depending on the thickness to be welded, the welding parameters used and possibly
the type of process employed. Useful indications must be submitted when the process is being
qualified, particularly with reference to the values of hardness obtained in the heat affected zone.
There is a slightly higher tendency to suffer lamellar tearing than with mild steels because the
ductility of steel in the shorter cross-section (in the direction of the thickness) is more negatively
influenced by small defects as its resistance grows.
From the point of view of brittleness ruptures, as these steels are always normalised and
manufactured with care, they can generally speaking be considered safe so long as the right type
has been chosen (from the various grades of toughness) in accordance with the service conditions,
bearing in mind residual voltages that are of higher values.
From this panorama on the weldability of low manganese alloy steels one can deduce that they
can be used in all processes applicable for mild steel, provided weld metal with a low hydrogen
content is used and pre-heating is at a higher temperature. Welding with covered electrodes is
almost always done using basic electrodes because of their low hydrogen content (if well-kept):
cellulose electrodes are used almost exclusively in welding methane and oil pipelines, for the
known technological reasons.
After welding, particularly in the case of constructions more subject to stress and for thicknesses in
excess of 25 mm, it is generally necessary to employ a stress relieving heat treatment.
111
High resistance quenched and tempered steels
Quenched and tempered steels are low alloy steels that reach extremely high yield point values
(from 500 to 900 N/mm2 – Fig. 198).
These steels owe their name to the fact that they are hardened during manufacture (generally
through high pressure water jets) and then brought back to a temperature of between 550 and
700°C depending on their chemical composition and the final mechanical properties required. The
composition is characterised by a low carbon content and the presence of other alloying elements
such as Cr, Ni, Mo, Cu as well as – of course – Mn and Si in amounts that vary according to the
type; small quantities of other elements like Nb, V, Ti and B might also be present in certain types.
Fig. 198
These steels are designed to achieve considerable resistance, thanks to their particular chemical
composition and heat treatment, and at the same time offer a satisfactory weldability naturally on
condition that the appropriate precautions have been taken.
They have a good grade of toughness for various reasons:
− the grain is very fine not just through the effects of normalising with Al, but also through the rapid
cooling from austenitic temperature to ambient temperature undergone during hardening
− thanks to the low carbon content, the martensitic steel thus obtained is relatively tough
− the contents of elements that prejudice the toughness (like S, P, Si, N) and alloying elements that
can influence it (such as V) must be controlled very carefully indeed, particularly for steels with
higher resistance.
112
Considering their resistance, quenched and tempered steels are relatively easy to weld with the
majority of welding processes, because toughness comparable to that of the base material can be
achieved without heat treatment.
Two fundamental rules have to be observed, however:
1. the cooling speed of the heat affected zone must be fast enough to allow the base material,
which exceeds the austenitic temperature during the welding process, to form a new
martensitic or bainite structure (otherwise resistance and toughness are compromised)
2. the quantity of hydrogen absorbed by the weld pool must be drastically limited (to avoid the
formation of hydrogen cracks in the fusion area and the heat affected zone).
If the first of the two rules is not observed, ferrite is produced in the heat affected zone (extremely
poor in carbon of course and therefore offering poor resistance) together with martensitic and
bainsitic areas (in which the carbon is concentrated and which are therefore not very tough), with
a consequential diminution of toughness and resistance.
The minimum cooling speed needed to achieve hardening varies according to the chemical
composition of the steel and depends mainly on the specific heat provided
Hi (J/cm) = (V x I x 60)/v
where: V = tension in volts; I = current intensity in A; v = productivity speed in cm/min
on the thickness of the edges and on the pre-heating temperature.
When welding these steels then, Hi must be lower than the maximum value established as a
function of the other two parameters to ensure a sufficiently severe thermal welding cycle.
Too severe a cycle can be dangerous however (to avoid excessively high points of hardness) and
moderate pre-heating is therefore advised for considerable thicknesses (above 15 – 30 mm
according to the type of steel).
With regard to the second of the two rules set out above, this appears obvious if we consider the
structure of the heat affected zone on one side and the hardness of the fused area which must be
adequate to guarantee the considerable resistance of the base material on the other.
In effect, not being able to avoid hard structures in the heat affected zone and the fusion area, for
these steels it is necessary to limit the quantity of hydrogen that might be absorbed by the weld
pool and adopt a light pre-heating so as to dry the edges and ward off the possible influence of
ambient humidity during the welding process.
Furthermore, pre-heating is generally very useful to facilitate evacuation of hydrogen that gets into
the weld pool.
We can observe, however, that more than the pre-heating it is important to eliminate the in situ
sources of hydrogen in certain processes. For example in this regard we can mention the coverings
of electrodes for manual welding and flux in submerged arc welding that always introduce a
certain quantity of hydrogen to the weld puddle.
In addition to the two fundamental rules set
out above, it is useful to employ a multiple
pass technique to benefit from the tempering
effect of subsequent passes on the previous
one and on their heat affected zones (Fig.
199).
Generally speaking this leads to the use in
welding of rectifying heat treatment on the
heat affected zone similar to that of
steelworks, in practice confining the
alterations mentioned above to limited areas,
or even eliminating them altogether if this can
be done through a precise and accurate
sequence of passes.
Fig. 199
At this point we can sum up the properties of weldability of these steels, noting that the greatest
danger comes from cold cracking, as we are normally dealing with hardened structures.
113
Hot cracking does not represent a big problem because these steels have low carbon and impurity
contents.
The same thing goes for lamellar tearing (very low content of sulphurs) although the poor ductility
could, under certain project conditions, lead to this risk.
There might be a certain risk of brittle ruptures in correspondence with service conditions
characterised by low temperatures and the presence of severe indentations, because the internal
stresses are very high for high resistance steel.
It might be necessary to join elements made of quenched and tempered steels to elements made
of carbon manganese steel. In this case the weld metal to be used can be one suitable for
welding quenched and tempered steel but preferably it should be more suitable to the least
resistant of the two metals being welded; obviously, in such cases the rules and precautions set out
above are to be followed scrupulously.
Molybdenum-chromium steels If mild steel is stressed beyond a certain limit at high temperature – above 350 °C – it progressively
lengthens or hot creeping takes place and is more rapid and intense as the temperature and
torsion increase.
This is due to the fact that the work hardening effect owing to the deformation, i.e. the increased
resistance deriving from distortion of the crystalline lattice, is gradually wiped out because, thanks
to atomic mobility that increases with temperature, a continual re-adjustment of the crystals takes
place as the atoms move.
Adding molybdenum (0,5 ÷ 1%) to the steel, the atoms in the molybdenum take the place of the
smaller iron atoms in the lattice, distorting the latter: this means that even at high temperature the
atoms find it harder to move around and the resistance to hot creep therefore increases. The
formation of extremely fine molybdenum carbides that are dispersed in the crystals contributes to
this by slowing down deformation, locking the creep planes in the crystalline lattice.
In a purely molybdenum steel, and above all in the heat affected zones of the weld, it may
happen that, as in carbon steel, the phenomenon of graphitisation may appear after prolonged
service at high temperature. Graphitisation is decomposition of the carbides with separation of the
graphite that makes the steel extremely fragile.
The addition of chromium to the steel (0,5 ÷ 5%) eliminates this problem (Fig. 200).
As the percentage of chromium increases, the steel’s chemical resistance to the oxidising gas and
sulphurs at high temperature increases too. However, when the temperature is extremely high, the
chemical resistance and resistance to hot creep in Cr-Mo steel becomes insufficient and it is
therefore necessary to use chromium or chrome-nickel stainless steel.
Fig. 200
If a carbon steel works at high temperature in the presence of high pressure hydrogen, the
hydrogen having smaller atoms will tend to spread to the steel and here react with the carbon to
form methane (CH4). Given the greater dimensions of the methane molecule, this cannot spread
and it therefore collects on the edges of the grains and on certain crystallographic surfaces
generating strong stresses, leading even to inter-crystalline fissures that make the steel extremely
fragile. To avoid this problem, chromium is added to the steel in quantity sufficient to enable stable
carbides to form, on which hydrogen cannot act.
The particular importance that chrome-molybdenum steels have assumed in producing boilers,
reactors, refinery piping, petrochemical and chemical plants etc derives from these overall
properties.
114
An exam of the TTT curves for Cr-Mo steels can show us how a completely ferrite-pearlite structure
(and therefore free of particularly hard or fragile formations) can be obtained only with very low
cooling speeds such as those that can be achieved in heat treatment but certainly not during
welding. Welding without pre-heating, and therefore with high cooling speeds, one will obtain in
the fusion zone and the heat affected zone a martensite structure, with a percentage of
martensite that increases as the cooling speed rises and the percentage of alloying elements
increases. Welding after proper pre-heating, one can obtain a prevalently bainite structure that
has considerably lower hardness and fragility, and therefore with lower risks of cracking. It should
also be remembered that the hardness of martensite obtained depends essentially on the
percentage of carbon: it is therefore necessary that the base metal, or better still the weld metal,
has a low carbon content.
The minimum pre-heating temperatures are 100 ÷ 150°C for steels with 0,5% Mo if the thickness is
greater than 15 mm and, according to the thicknesses, 100 ÷ 200°C for steels with 0,5% Cr – 0,5%
Mo, 150 ÷ 250°C for carbon content of between 0,8 and 1,25% and 200 ÷ 250°C for 2,25% Cr – 1%
Mo up to 9% Cr – 1% Mo steels. These minimum temperatures must of course be maintained
throughout the weld. It is also wise to heat the entire joint uniformly to a temperature higher than
that of pre-heating once welding has been completed, leaving it then to cool down slowly or
proceed with heat treatment immediately after welding has finished.
Because excessive hardness grades are obtained despite welding with pre-heating, it is necessary
to proceed with tempering and normalising heat treatment after welding at a temperature slightly
lower than the Ac1 temperature or annealing at a temperature slightly higher than Ac3; in the latter
case, cooling must necessarily be slower so as not to obtain hardened structures once again.
From the transformation diagrams one can see how the formation of chromium carbides takes
place only with very low cooling speeds: this observation assumes particular importance for Cr-Mo
steels that have to resist hydrogen attacks, given that the formation of chromium carbides is
essential to that end. With a tempering heat treatment a tempering structure will be obtained in
which the chromium carbides will have precipitated while, with an annealing heat treatment and
very slow cooling, a structure with ferrite, pearlite and carbides will be obtained.
Tempering and normalising heat treatment will be carried out at 620 ÷670°C for 0,5% Mo steel (for
this steel and thicknesses of less than 20 mm the heat treatment can be omitted), at 630 ÷680°C for
0,5% Cr – 0,5% Mo steel, at 660 ÷710°C for 0,8% Cr – 0,5% Mo and 1,25% Cr – 0,5% Mo, at 730 ÷ 780°C
for 2,25% Cr – 1% Mo and at 700 ÷ 750°C for 5% Cr – 0,5% Mo.
Heat treatment times are generally set as 1 hour for each inch of thickness with a minimum of 30
minutes.
We have seen how these steels tend to develop cold cracking and the possible remedies; hot cracking is generally not a problem because we are dealing here with high quality steels with
extremely low rates of impurity; given the general use of limited thicknesses and the low
percentages of impurities, lamellar tearing is not one of Cr – Mo steel’s typical defects either.
The welding processes generally used are: TIG (first passes) and manual basic covered electrode
welding; at times, but not very often, MIG/MAG processes can be used. The automatic submerged
arc process employing a combination of suitable fluxes and wires can be used if the heat
conveyance is contained.
Nickel steels The first and most economical metallurgical practice for improving the toughness of the most
common C-Mn construction steels is that of a final deoxidisation through the addition of small
quantities of aluminium, which brings about the following advantages for semi-worked pieces
generally supplied in the normalised state:
− complete stress relief: the consequential excellent deoxidisation and the uniform distribution of
the residual impurities throughout the mass is the preferred situation in which to start welding,
avoiding local concentrations caused by segregation and improving the steel’s ductility and
toughness
− a fine grain structure that is always very favourable both in terms of mechanical resistance and
above all toughness
− locking of residual nitrogen in hardly soluble elements; the steel becomes anti-ageing.
115
With this practice, common C-Mn constructions steels can safely be used at service temperatures
of -30 and even -40°C (Fig. 201).
But the perfect deoxidisation and the ferrite-pearlite fine grain structure do not allow much more.
This is where the addition of an alloying element – nickel – comes into play and special steels are
produced that are suitable also for welding. The metallurgical properties of these steels are due
mainly to the nickel and the resulting mechanical properties depend on the heat treatment that
the semi-manufactured parts undergo after being hot-worked.
The metallurgical function of nickel added to iron manifests itself through remarkable austenitising
properties.
Nickel is easily soluble in γ iron and α iron and reinforces the crystalline lattice of the iron, increasing
in practice the resistance of the steel without increasing its hardness too much. But the most
interesting effect consists in a large and general lowering of the allotropic transformation
temperatures of the iron (A3, A1, Ms, Mf) and a large decrease in the critical tempering speed. For
these reasons, adding appropriate amounts of nickel will make the steel all the more sensitive to
heat treatment, a very fine grain will be obtained, and hardening will penetrate in depth allowing
the manufacture of semi-worked products with a good structural homogeneity throughout the
thickness, both on the surface and at the heart of the piece. Basically, the presence of nickel in
the steel has a strong influence on the structures that can be obtained by heat treatment and, by
adding the appropriate amount of this element, the desired structures – pearlite, martensite,
austenite – can be achieved.
Our welding notes below will deal with two of the various types of nickel steels: 3,5% Ni and 9% Ni;
2,5% Ni differs from 3,5% Ni in welding only in that, with the former, lower alloy weld metals are used
in relation to the service temperature, generally between -50 ÷ 60°C for 2,5% Ni base metal. From a
welding point of view, the 5% Ni base metal does not present any real construction advantages
with respect to 9% Ni, which is more costly but certainly more reliable in its resilience at
temperatures in the order of -150°C or lower.
Fig. 201
3,5% Ni steel has minimum guaranteed mechanical resistance only slightly lower than that of C-Mn
type Fe 510 steel. It can be supplied at various heat treated stages: normalised, normalised and
tempered, or quenched and tempered, and the metallographic structure that is presented can
vary in relation to the content in carbon and the heat treatment undergone, from ferrite/pearlite to
very fine grain to partially or totally bainite.
The micro structure and the resulting toughness of the steel are particularly sensitive to tempering
temperatures: although they have a theoretical Ac1 transformation point of 670°C, it can
sometimes happen that the original toughness of the material is negatively altered even tempering
116
at 620°C through a probable separation of richer austenite into carbon and nickel that can be
transformed yet again in the next cooling, leading to more fragile structures.
We are dealing with a steel that can suffer brittleness during tempering and in such cases the need
to control the welding and heat treatment temperatures with more care than would normally be
suggested for handling common C-Mn steels is well known. Then again, because this steel has a
greater tendency to harden than C-Mn steels, and because it is in any case better to reduce the
residual welding stresses as much as possible for low temperature service, it is common practice to
weld after a light pre-heat (100 – 120°C) at least where the thickness is greater than 10 mm, then
performing the tempering and normalising heat treatment at a temperature of 550-610°C.
Generally speaking it is a good rule not to perform heat treatment at temperatures greater than
that undergone by the base metal during manufacturing of the semi-worked piece. In particular
cases it can be preferable to leave out heat treatment after welding, taking into account the
effective service conditions of the welded construction.
With or without heat treatment, the practice of controlling the weld heat limiting the temperature
between passes is enough to keep toughness in the heat affected zone to within acceptable limits,
even though at the same time it may drop slightly with respect to the original base material.
Less predictable results might be found in the fusion zone of the weld when the service
temperatures are at the limits of those for which the base material was designed (currently down to
about -90°C, liquid ethylene). Designing good weld materials with composition similar to the base
metal has been a very delicate matter because of the hot cracking phenomenon that is more
accentuated than for C-Mn steels; the analysis elements C, Mn, Si must be controlled to within
much tighter limits than for the base material and the harmful influence of the impurities (S, P) is
more marked. However, with modern weld metals, taking care to ensure utmost cleanliness and by
limiting the specific welding heat (moderate pre-heating, controlling the temperature in between
passes, not too high a current), the phenomenon of hot cracking in the fused zone is generally
limited, when it does exist, to the craters at the end of the pass, so it is a good rule to eliminate
these with a grinder in between each weld pass.
The weld zone’s sensitivity to tempering temperature and time is a phenomenon not yet controlled
properly. In practice, a scattering of results is found due to the toughness in the weld zone varying
from case to case, and it is widely held that understanding the chemical analysis of the weld
puddle is not sufficient to guarantee the results beforehand, at least when the service temperature
is going to be low.
If possible, welding positions other than flat are avoided, both because of the known drop in
resilience that vertical welding leads to and because of the importance given to the absence of
internal defects and even external welding defects (surface irregularities, cuts) for low temperature
service.
Compared to 3,5% Ni, 9% Ni with regard to its weldability in a broader sense must take into account
some further factors:
- this is a very high resistance steel because 9% Ni has a tempered martensite structure. There is
therefore the problem of achieving similar mechanical resistance in the joints.
- due to the accentuated phenomenon of cracking and poor toughness in the weld zone, no
weld metals with a composition similar to that of the base material have yet been invented.
Current solutions include using austenitic alloys as the weld material. These are costly and can
hardly offer the same toughness in the weld area as the base metal while they do afford good
performance from the point of view of the toughness at low temperature. The calculus of the
joint’s resistance is therefore done on the properties of the fused area, which is weaker, and from
this point of view the current situation can be considered a way around the problem however
refined and consolidated the technique may be.
9% Ni steel is supplied in two states of heat treatment: with double normalising and tempering
(normalising at 900°C + normalising at 790°C + tempering at 565 ÷ 605°C and cooling at a rate of
>165°C/hour) or rectifying (hardening at 800°C in water + tempering at 565 ÷ 605°C at a speed of
>165°C/hour).
The metallurgical structure of the steel is tempered martensite with a residual percentage of
austenite around 5÷20% (according to the cycle of heat treatment).
Because of the high content of nickel, this steel presents a characteristic viscosity at structural
transformations and is very sensitive to the heat history it undergoes in the interval between
300÷700°C; this thermal history obviously refers to the heat affected zone the structure of which
cannot avoid being disturbed by the thermal cycle of welding. The considerations we have put
117
forward concerning the heat affected zone in quenched and tempered steels are also very valid
here for 9% Ni steel too, as this is also a high resistance treated steel. Welding of this steel requires
controlled heat conveyance during the weld passes to limit damage caused by this heat in the
heat affected zone.
Being a more or less tempered martensite, the hardness met with in the heat affected zone will be
higher than that found in a high resistance C-Mn steel and the construction specifications in fact
allow maximum values of around 400 HV10. Then again, it should be remembered that the
hardness of the base material unaltered by welding is already around 230÷260 HV10.
The use of austenitic weld metal allows post-welding heat treatment to be avoided (something this
material is very delicate in relation to) and the known risks connected with hydrogen’s effect on the
heat affected zone also diminish.
The weld metals used with these steels must therefore deposit an austenitic type structure even
after dilution in the steel base metal. To this end alloys with high nickel content (such as INCONEL)
are used or even 309 or 316 stainless steels with the addition of other austenitising elements and
elements that reinforce the matrix like manganese, tungsten or cobalt, each in percentages that
go from 3 ÷ 8%.
The purpose of these additions of reinforcing elements is to increase the mechanical resistance in
the weld zone to match that of the base metal.
Defects that might be picked up in the weld zone are hot cracks, pertinent to the austenitic
structure: the risk of this defect usually involves the first pass done quickly and with considerable
dilution of the base material using an automatic process (MIG or submerged arc).
However, precautions such as cleanliness, observance of the temperature limits between
subsequent passes, not too large a weld puddle, suitably limited electrode diameter and current
intensity, and removal of craters at the end of each pass allow this aspect to the minimised even
for 9% Ni steels.
Austenitic chromium-nickel stainless steels Compared to pure chromium stainless steels, Cr-Ni alloy steels present some substantial
improvements with regard to the mechanical resistance at high temperature (hot creep) and their
workability in general, with particular emphasis on the weldability, while having anti-corrosion
properties that are generally equal to those of chromium steel, except in a reductant or sulphur
contaminated ambient as this attacks the nickel.
The chemical composition of the main types used in welded constructions is characterised by the
contents of chromium (18 ÷ 25%) and nickel (8 ÷ 20%), which are the two basic constituents of this
range of steels and proportioned in such as way as to obtain a final structure that is totally or mainly
austenite (Fig. 202).
Fig. 202
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The austenitic structure confers excellent anti-corrosion and mechanical properties on the alloy as
well as excellent workability, which might suggest using it in place of chromium ferrite steels, despite
the higher cost due to the presence of nickel.
The first austenitic stainless steel to be made was the 18-8 type, which represents the lowest content
in the principal alloying elements – chromium and nickel – needed to achieve the austenitic
structure at ambient temperature.
Other elements in addition to chromium and nickel, added in smaller proportions, improve
determined properties of these alloys: molybdenum confers greater resistance to corrosion in a
reductant ambient and, above all, it improves the mechanical resistance at high temperature;
generators of carbides such as titanium and niobium, apart from contributing to improving the hot
mechanical resistance, are added to avoid a drop in the anticorrosion properties.
Compared to chromium ferrite steels, it should also be noted that austenitic chromium-nickel steels
do not present a rapid drop in their ductility related properties (transition) even at the lowest of
temperatures required, for example, in the liquid gas industry, for which they are therefore well
suited despite their higher cost.
Nickel is a metal that melts at 1445°C and crystallises in a central faced cubic lattice, as ferro-
gamma. It does not present any allotropic transformations. Added to iron, it has an austenitising
effect widening the field in which the austenitic phase takes place.
Added to carbon steel or to chromium and nickel, it is considered a hardening element because it
lowers the critical hardening speed.
18% Cr ferrite steel with the addition of 2% Ni becomes susceptible to martensitic hardening as 13%
Cr steel. Increasing the nickel content will widen the zone in which the austenitic phase takes
place and progressively reduce the ferrite + carbide zone. The addition of 4% Ni will result in a
hardenable steel starting at 900°C but already austenitic through hardening at 1100°C; the
austenite thus obtained at ambient temperature is unstable, however, and might be transformed
following tempering.
With 8% Ni the critical hyper-temper speed becomes low enough to achieve austenite at ambient
temperature without the gamma-alpha transformation and precipitation of the carbides taking
place; the carbon remains in a supersaturated state in the austenitic matrix and the steel presents
a homogeneous structure with maximum resistance to corrosion.
Austenitic Cr-Ni steels are therefore used, in normal conditions, at the hyper-temper state that is
reached through heat treatment that makes the carbides in the austenitic soluble (generally
T=1050°C, i.e. to be within the stable gamma field) and subsequent rapid cooling.
The proportion between the contents of chromium and nickel and the other austenitising elements
(such as carbon and manganese) and ferritising elements (such as molybdenum, silica, niobium)
that might be present or added on purpose is usually chosen so as to achieve a structure that is still
austenitic but with increasing chromium content that confers the basic inoxidisable property on the
alloy.
In addition to stabilising the austenitic phase and improving corrosion resistance, nickel acts
favourably also as a refining element of the structure which, in theory, not having a gamma-alpha
transformation line, should always present the problem of a visible swelling of the grain through
dwelling at high temperature.
The swelling takes place in the heat affected zone of austenitic welds but does not reach the
proportion and level of danger that one has in chromium ferrite stainless steel welded joints.
For chromium-nickel steels there are two known structural modifications caused by dwelling at
temperatures of around 500÷800°C, which are particularly serious for two reasons: they deteriorate
the most characteristic properties of austenitic steels (high toughness and corrosion resistance) and
this happens in a field of temperatures for which those properties stand out most of all with respect
to lower alloy steels.
Precipitation of chromium carbides
As we have said, to keep the carbon in the austenite in a supersaturated solution, stainless steels
are supplied in the hyper-temper state. Obviously a subsequent re-heating at 450÷750°C will favour
a return to equilibrium with a precipitation of the chromium carbides, The carbides arrange
themselves around the austenitic grains and compromise the corrosion resistance; to feed the
carbides, the grains’ surroundings are poor in chrome and reduced corrosion resistance is
119
attributed to this fact (Fig. 203 Precipitation of chromium carbides at the edge of the grain – 204
consequent inter-granular corrosion – SEM images).
Fig. 203 Fig. 204
The phenomenon is influenced by the content in carbon, the temperature and the dwell time; the
lower the level of carbon, the slower corrosion will be after precipitation; the lower the precipitation
temperature, the slower the precipitation: with C=0,08% 20 seconds at 700°C is enough, with
C=0,03% 10 hours at 600°C is needed.
Bringing the material back to above the precipitation interval, the carbides return to the solution.
The weld zone does not fear the precipitation because it exceeds the critical interval to re-cross
that line very quickly again during cooling.
The main remedies against carbide precipitation are:
- solubilisation heat treatment (heating at T > 1000°C and rapid cooling), which as we saw puts
the carbides back into the solution. This treatment is not used much on finished constructions
because of its high cost and the deformations it causes.
- decarburisation or the production of very low carbon content steels; experimentally it has been
found that the effects of precipitation are not a cause for concern when the carbon content is
lower than 0,03% (maximum limit for L type steels). The low carbon content leads to a certain
reduction in the resistance to traction.
- stabilisation, that is the addition of alloying elements, referred to also as stabilisers, capable of
forming their own carbides quicker than chromium does. These elements are titanium, niobium
and tantalum. The carbides of these elements remain uniformly distributed in the austenitic
matrix (Fig. 208) and, differently to chromium carbides, they do not form continuous chains
around the grains. The chromium remains undisturbed and the steel conserves its corrosion
resistance in full.
When the stabilising agent is tantalum the steel must contain at least five times the amount of
carbon whereas when niobium and titanium are used their sum must add up to at least 10 times
the quantity of carbon. The most well-known stabilised steels are AISI 321 and AISI 347. When
welding stabilised steels the right choice of weld metal is one that deposits pure stabilised weld
pools; in this regard it should be remembered that stabilised type covered electrodes all use
niobium and tantalum because titanium evaporates in the arc and therefore cannot be
transferred.
Stabilised steels can risk the following scenario: the Ti, Nb and Ta carbides form rather slowly in the
interval between 900 and 1200°C; they return to the solution very quickly above 1200°C. In
preparing stabilised sheets the steelworks will arrive at 1100°C, dwelling for the time it takes for
carbides to form, and then it is cooled rapidly. During welding, the temperature of the thermal
cycles exceeds 1200°C and the stabilised carbides are destroyed, returning to the solution in a thin
layer of base material adjacent to the bead. During cooling, the precipitation interval (1200-
900°C) is crossed very quickly, the carbides do not reform and the carbon remains in a
supersaturated solution up until ambient temperature. If the construction is taken to a service
temperature of between 500 and 800°C, the thin strip of metal adjacent to the weld bead
becomes the place where carbides collect in precipitation. The phenomenon may manifest itself
immediately after welding when, to avoid corrosion under stress, the piece is taken through
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normalising heat treatment at 700°C. It is thus shown that, despite stabilisation, precipitation of
chromium carbides can take place with consequential corrosion (called knife blade).
The phenomenon is extremely rare and manifests itself only in the presence of energetic aggressive
chemicals such as nitric acid.
The sigma phase
Intermetallic Fe-Cr-Ni compost susceptible to form between 550 and 900°C is called sigma phase.
This compost is hard and brittle and precipitates in needle shapes inside the grains (Fig. 205 & 206),
influencing the mechanical properties and the corrosion resistance.
Both carbides and the sigma phase can be observed on a microscope with a x1000 magnification.
To destroy the sigma phase it can be put back into the solution heating the metal to beyond 950°C
followed by cooling at a rate greater than 100°C/hour.
Because precipitation takes place in slow cooling, heat cycles of welding are too rapid to form the
sigma phase, which is therefore something to be concerned about only as a consequence of heat
treatment or prolonged service at high temperature.
Fig. 205 Fig. 206
Thanks to their austenitic structure, stainless steels do not fear cold cracking; on the contrary, hot cracking might take place during the initial solidifying and cooling stages.
As for common steels, austenitic stainless steels may suffer from hot cracking caused by
segregation of the sulphur and phosphor which, forming a liquid veil over the edges of the grains,
weakens the section of the weld zone called upon to face the internal stresses; niobium
segregation can also take place, which is why niobium stabilised steels crack more easily. Silica
can also segregate. We would also mention carbon, noting that low carbon steels are less likely to
crack.
Stainless steels have a higher expansion coefficient and lower thermal conductivity than common
steels. The heat cycles of welding therefore lead to expansion that is more pronounced in the less
extended regions, reinforcing the internal stresses that are always a determining factor in hot
cracking.
Another unfavourable circumstance lies in the fact that the dimensions of the austenitic grains are
on average larger than those of carbon steel so they are more sensitive to segregation.
The presence of ferrite in the austenitic matrix (Fig. 207 & 208) of the weld zone provokes both
negative and positive consequences. First of all we should remember the favourable effect ferrite
has against chromium carbides. When there is ferrite present, the carbides prefer to deposit inside
the ferrite grains rather than around the austenitic grains; because the chromium content is higher
in the ferrite than it is in the austenitic grains, the consequential loss of chromium worries us less than
it would in mono-phase steels.
Ferrite is also a precious antidote against hot cracks because it slows down swelling of the
austenitic grains, increasing considerably the total surface area surrounding the grains.
Ferrite is not convenient in the sigma phase because it accelerates precipitation.
Ferrite can compromise the corrosion resistance of austenitic stainless steel if other chemical agents
are present, such as might happen in plants for the production of urea for example. The
phenomenon is attributed to the difference in the compositions of the ferrite and the austenite as
121
well as to the diminution of chromium around the edges of the austenitic grains caused by the
formation of bands of ferrite that are richer in chromium.
Fig. 207 Fig. 208
To estimate the percentage of ferrite on the basis of the composition of the weld puddle, we can
make use of Schaeffler’s diagram (Fig. 209).
Fig. 209
Both Schaeffler’s diagram and the formula serve only to give a rough approximation seeing that,
with identical compositions, the ferrite content can vary with the cooling speed and might be
modified by heat treatment.
There are several instruments available in commerce for measuring the ferrite in weld beads. They
are based on the fact that, differently to austenite, ferrite is magnetic; certain instruments measure
the permeability, others the magnetic saturation.
Some orders of size for ferrite: 8% ferrite is enough to form a continuous grid around the austenitic
grains, 4% ferrite is the lowest recommended content to limit crack tendency in stabilised steels; 3%
ferrite is the maximum content allowed nowadays in urea plants.
Particular points that need bearing in mind when welding are:
- quick and thin passes to accelerate setting and cooling time
- inter-pass temperature limited to 100-150°C
- particular attention to be paid to craters at the end of the bead
122
- in submerged arc welding, characterised by high levels of heating, it should be taken into
account that there are more voluminous passes which, while cooling slowly, maintain the weld
zone inside the interval of hot brittleness for longer.
In cases of welding dissimilar steels (carbon steels to austenitic stainless steels) the aspect that has
to be evaluated most carefully is that of dilution of the austenitic weld material in the ferrite base
metal; in particular cases of dilution the formation of hard fragile (martensite or similar) structures
may happen in the weld area (generally around the fusion line on the carbon steel side) with an
obvious risk of cracking.
Predicting the structure with good approximation can be done using Schaeffler’s diagram; during
the process qualification stage, a careful check on the hardness and ductility values and those of
the bending test can be very significant.
Generally, this type of joint is done using “highly austenitic” weld metal type 309 or 310, preferably
with the addition of molybdenum.
Austeno-ferritic stainless steels (DUPLEX) Duplex stainless steels with a twin structure, austenite and ferrite, and with a high carbon content
have been known for some time: compared to austenitic steels, these have better corrosion
resistance in a chlorinated environment. From a welding point of view, as already mentioned for
austenitic stainless steels with small percentages of ferrite, these are sensitive to inter-granular
corrosion and through their tendency to form rough structures they result as being rather fragile in
their heat affected zone.
These problems led to the development of austeno-ferritic steels with low carbon levels that have
excellent resistance to tension-corrosion and inter-granular corrosion and good mechanical
properties. These steels also offer resistance to pitting in seawater or a chlorinated environment.
Duplex steels contain chromium in amounts of between 18,5 and 27%, they have between 4 and
8% nickel and between 2 and 4% molybdenum (Fig. 202); their carbon content is lower than 0,03%;
they are characterised by a microstructure that contains both austenite and ferrite. The most
widely used type contains 22% chromium and 5% nickel (type 2205) and has a structure that is
approximately 50% ferrite and 50% austenite (Fig. 210-211). The favourable properties of these
steels can be attributed to the optimal ratio between ferrite and austenite, achieved from a
correct equilibrium between the ferritising elements (Cr-Mo) and the austenitising elements (Ni-Mn-
N).
Fig. 210: rolled duplex Fig. 211: cast duplex
The weldability of duplex steels is controlled by the structural transformations that take place in the
heat affected zone and the fusion zone. In the heat affected zone, in which the temperature
exceeds 1200°C, the austenite is re-transformed into ferrite-delta and, after the subsequent cooling,
ferrite might be present in percentages higher than the initial levels. In the heat affected area
though, where the temperature does not exceed 1000°C, a part of the ferrite is transformed into
123
austenite during the heating stage; with the successive rapid cooling, the gamma phase remains
stable: there is therefore a diminution in the final count of ferrite-delta content.
With regard to the weld pool, the equilibrium between ferritising and austenitising elements
contained in the weld metal is important; in practice, the resulting quantity of ferrite in the weld
pool must be calculated so as to get as close as possible to that of the unaltered base metal. One
must also bear in mind that heat cycles that are too bland could lead to an excessive swelling of
the grain and precipitation of chromium nitrides with a reduction in the corrosion resistance, while
overly severe heat cycles could excessively increase the percentage of ferrite delta with a
consequential deterioration, in particular, of the ductility and toughness properties.
An important effect in these steels is
produced by the presence of nitrogen (Fig.
212): in an N=0,05% steel there may be ferrite
values in excess of 80%; the addition of 0,15%
N induces ferrite reduction: the start of the
transformation from austenite to ferrite is
higher at higher temperatures: it follows that
there is a greater quantity of austenite than
there is of ferrite. The greatest effect is that
found in steel with high nitrogen and nickel
content.
One other favourable effect of nitrogen
derives from the fact that in the weld pool, as
in the heat affected zone close to the line of
fusion, the formation of rough ferritic grains is
stopped by the presence of nitrogen itself.
The weld metals used generally have a
chemical analysis similar to that of the base
metal; for the already-mentioned 2205 the
average analysis of the deposition weld
metal used is 22% Cr + 9% Ni (2209).
Fig. 212
Aluminium and aluminium alloys
The mechanical properties of pure aluminium are insufficient for many applications (Rm around 70-
80 N/mm2 in the annealed condition).
Hence the creation of many aluminium alloys, known as light alloys on account of their low density
(approximately 3 g/cm3).
The temperatures, or rather the ranges of fusion, are between 500 and 660°C, depending on the
composition.
The most commonly used alloying elements are: Mg, Si, Mn, Cu, Zn, Fe, Cr, either separately or in
combination (Fig. 213).
The great majority of aluminium alloys are highly suited to plastic working, namely the production of
plates, pipes, section bars and wires. The alloying elements are soluble up to certain levels of
concentration beyond which they form intermetallic compounds.
It follows that many alloys can be heat treated; in this respect we can distinguish between:
- alloys that can be quenched and tempered or strain hardened (5000 series alloys)
- alloys that can be quenched and tempered (6000 and 7000 series alloys)
Obviously the welding heat cycles will have the effect of upsetting every strain hardened or
quenched and tempered structure in the heat affected zone.
124
Fig. 213
Mention has already been made, with reference to the TIG welding process, of the problems
caused by the film of oxide that forms on the surfaces of these alloys; Al2O3, also known as alumina,
has the following properties:
- it melts at over 2500°C and therefore has excellent refractory characteristics
- it has greater specific weight than metal
- it is very stable and only strong bases (sodium carbonate, potassium carbonate) are able to
attack it.
It is evident that alumina presents many welding problems and in fact the metal reaches fusion
while the surface film of oxide is still in the solid state.
Therefore, the welder runs the risk of “burn through” before noticing that fusion of the metal has
occurred. The situation is made even more critical by the fact that aluminium does not change
colour when it heats.
The oxide wraps around the droplets of hot metal, preventing their union; if, for example, we take
two aluminium rods and we try to weld their ends together using an oxyacetylene torch we find
that we are unable to do so.
The surface film of oxide remains in the solid state and the droplets do not join.
The oxide also tends to form inclusions in the weld pool because it has greater density than metal.
The oxide inclusions seriously impair all the mechanical properties insofar as they disrupt metallic
continuity. Welding is only possible if the oxide is completely removed.
We have already seen how this is possible using the TIG process, with electric supply from a source
of alternating current.
With the MIG process, using a pulsed-arc, the results are excellent with significantly higher
productivity than the TIG process.
The other welding processes are not used for aluminium and aluminium alloys.
Another very important factor is cleaning, both of the joint and of the filler material (stick wire or
wire), largely so as to avoid the problem of blowholes.
In welding alloys that can be strain hardened (types H111/H112/H116/H321), the effect of strain
hardening disappears entirely so during the design stage (and as regards the reference values for
the tensile tests at the time of qualification of the welding process), irrespective of the supply
condition of the alloy, the mechanical characteristics of the metal that are to be taken into
consideration are those in the annealed condition; the filler material will accordingly need to be
chosen so as to satisfy these characteristics.
In alloys that can be quenched and tempered (types T5/T6), the effect of such treatment will last
for varying lengths of time depending on the composition of the alloy and the speed of the heat
cycles.
