Welding Technology
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ISF – Welding Institute RWTH – Aachen University
Lecture Notes
Welding Technology 1 Welding and Cutting Technologies
Prof. Dr.–Ing. U. Dilthey
Table of Contents Chapter
Subject
Page
0.
Introduction
1
1.
Gas Welding
3
2.
Manual Metal Arc Welding
13
3.
Submerged Arc Welding
26
4.
TIG Welding and Plasma Arc Welding
43
5.
Gas– Shielded Metal Arc Welding
56
6.
Narrow Gap Welding, Electrogas - and Electroslag Welding
73
7.
Pressure Welding
85
8.
Resistance Spot Welding, Resistance Projection Welding and Resistance Seam Welding
101
Electron Beam Welding
115
10.
Laser Beam Welding
129
11.
Surfacing and Shape Welding
146
12.
Thermal Cutting
160
13.
Special Processes
175
14.
Mechanisation and Welding Fixtures
187
15.
Welding Robots
200
16.
Sensors
208
Literature
218
9.
0. Introduction
2003
0. Introduction
1
Welding fabrication processes are classified in accordance with the German Standards DIN 8580 and DIN 8595 in main group 4 “Joining”, group 4.6 “Joining by Welding”, Figure 0.1.
2 Forming
1 Casting
4.1 Joining by composition
4.2 Joining by filling
3 Cutting
4.3 Joining by pressing
4.4 Joining by casting
4 Joining
4.5 Joining by forming
4.6.1 Pressure welding
5 Coating
4.6 Joining by welding
4.7 Joining by soldering
6 Changing of materials properties
4.8 Joining by adhesive bonding
4.6.2 Fusion welding
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Production Processes acc. to DIN 8580
Figure 0.1
Welding: permanent, positive joining method. The course of the strain lines is almost ideal. Welded joints
Screwing
show therefore higher strength properties than the joint types depicted in Figure 0.2. This is of advantage,
Riveting
especially in the case of dynamic stress, as the notch effects are lower.
Adhesive bonding
Soldering
Welding
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© ISF 2002
Connection Types
Figure 0.2
0. Introduction
2
Figures 0.3 and 0.4 show the further subdivision of the different welding methods according to DIN 1910.
Production processes 4 Joining 4.6 Joining by welding
4.6.1 Pressure welding
4.6.2 Fusion welding
4.6.1.1 Welding by solid bodies
4.6.1.2 Welding by liquids
4.6.1.3 Welding by gas
4.6.1.4 Welding by electrical gas discharge
4.6.1.6 Welding by motion
4.6.1.7 Welding by electric current
Heated tool welding
Flow welding
Gas pressure-/ roll-/ forge-/ diffusion welding
Arc pressure welding
Cold pressure-/ shock-/ friction-/ ultrasonic welding
Resistance pressure welding © ISF 2002
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Joining by Welding acc. to DIN 1910 Pressure Welding
Figure 0.3
Production processes 4 Joining 4.6 Joining by welding
4.6.1 Pressure welding
4.6.2 Fusion welding
4.6.2.2 Welding by liquids
4.6.2.3 Welding by gas
4.6.2.4 Welding by electrical gas discharge
4.6.2.5 Welding by beam
4.6.2.7 Welding by electric current
Cast welding
Gas welding
Arc welding
Beam welding
Resistance welding
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Joining by Welding acc. to DIN 1910 Fusion Welding
Figure 0.4
1. Gas Welding
2003
1. Gas Welding
3 Although the oxy-acetylene process has been introduced long time ago it
3
is still applied for its flexibility and mo4
6
5 8
7 9
1 2
bility. Equipment for oxyacetylene welding consists of just a few elements, the energy necessary for welding can be transported in cylinders, Figure 1.1.
1 2 3 4 5 6 7 8 9
oxygen cylinder with pressure reducer acetylene cylinder with pressure reducer oxygen hose acetylene hose welding torch welding rod workpiece welding nozzle welding flame
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Figure 1.1 3
density in normal state [kg/m ]
oxygen
propane
1.43 0.9
ignition temperature [OC] 600
ral gas; here C3H8 has the highest
400
calorific value. The highest flame in-
200
tensity from point of view of calorific
0
645
645
value and flame propagation speed is, 3200
flame temperature with O2 flame efficiency with O 2 flame velocity with O2 43 1350
2850 2770 0
300
490 335
510 natural gas
C2H2, lighting gas, H2, C3H8 and natu-
however, obtained with C2H2.
1.17
propane
1.2. Suitable combustible gases are
1.29 air
oxygen and a combustible gas, Figure
2.0
air
exothermal chemical reaction between
2.5 2.0 1.5 1.0 0.5 0
oxygen
Process energy is obtained from the
°C
10.3
370
8.5
330
KW k
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Figure 1.2
/cm2
cm
/s © ISF 2002
1. Gas Welding
4 C2H2 is produced in acetylene gas
loading funnel
generators by the exothermal transformation of calcium carbide with wa-
material lock
ter, Figure 1.3. Carbide is obtained from the reaction of lime and carbon in the arc furnace. gas exit feed wheel
C2H2 tends to decompose already at a pressure of 0.2 MPa. Nonetheless, grille
commercial quantities can be stored
sludge
when C2H2 is dissolved in acetone (1 l of acetone dissolves approx. 24 l of C2H2 at 0.1 MPa), Figure 1.4.
to sludge pit br-er1-03.cdr
© ISF 2002
Acetylene Generator
Figure 1.3 Acetone disintegrates at a pressure of
acetone
acetylene
more than 1.8 MPa, i.e., with a filling pressure of 1.5 MPa the storage of 6m³ of C2H2 is possible in a standard cylinporous mass
der (40 l). For gas exchange (storage and drawing of quantities up to 700 l/h)
N
a larger surface is necessary, therefore
acetylene cylinder acetone quantity :
~13 l
the gas cylinders are filled with a po-
acetylene quantity :
6000 l
rous mass (diatomite). Gas consump-
cylinder pressure :
15 bar
tion during welding can be observed from the weight reduction of the gas filling quantity : up to 700 l/h
cylinder. br-er1-04.cdr
© ISF 2002
Storage of Acetylene
Figure 1.4
1. Gas Welding
5 Oxygen duced
gaseous
is by
profrac-
cooling
tional distillation
cylinder
nitrogen air
of liquid air and bundle
stored in cylinders
oxygen
liquid air
with a filling pres-
pipeline liquid
oxygen
sure of up to 20 MPa, Figure 1.5.
tank car nitrogen vaporized cleaning
compressor
For higher oxygen
separation
consumption, stor-
supply
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age in a liquid state
© ISF 2002
Principle of Oxygen Extraction
and cold gasification is more profit-
Figure 1.5
able.
The standard cylinder (40 l) contains,
50 l oxygen cylinder
at a filling pressure of 15 MPa, 6m³ of
protective cap cylinder valve
O2 (pressureless state), Figure 1.6.
gaseous
take-off connection
N
Moreover, cylinders with contents of
p = cylinder pressure : 200 bar
10 or 20 l (15 MPa) as well as 50 l at
V = volume of cylinder : 50 l Q = volume of oxygen : 10 000 l
20 MPa are common. Gas consumpcontent control
tion can be calculated from the pres-
Q=pV
sure difference by means of the gen-
foot ring
eral gas equation. manometer
liquid
safety valve
vaporizer
filling connection user
still liquid br-er1-06.cdr
Storage of Oxygen
Figure 1.6
gaseous
1. Gas Welding
6
In order to prevent mistakes, the gas cylinders are colour-coded. Figure 1.7 shows a survey of the present colour code and the future colour code which is in accordance with DIN EN 1089. The cylinder valves are also of show a thread
right-hand union
Acetylene
different designs. Oxygen cylinder connections
actual condition
nut.
DIN EN 1089
blue
actual condition
white
DIN EN 1089
grey
cylinder
helium
oxygen techn.
valves are equipped
yellow
brown grey
blue (grey)
brown
red
dark green
grey
with screw clamp acetylene
retentions. Cylinder valves
for
grey
other argon
a
darkgreen
left-hand
vivid green grey
grey
combustible gases have
hydrogen
argon-carbon-dioxide mixture black
grey
grey
darkgreen
thread-connection
nitrogen
carbon-dioxide
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with a circumferen-
© ISF 2002
Gas Cylinder-Identification according to DIN EN 1089
tial groove. Figure 1.7
Pressure regulators reduce the cylinder pressure to the requested working pressure, Figures 1.8 and 1.9. cylinder pressure
working pressure
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Single Pressure Reducing Valve during Gas Discharge Operation
Figure 1.8
1. Gas Welding
7
At a low cylinder pressure (e.g. acetylene cylinder) and low pressure fluctuations, single-stage regulators are applied; at higher cylinder pressures normally two-stage pressure regulators are discharge pressure
locking pressure
used. The
requested
pressure is set by the
adjusting
screw. If the pressure increases on the low pressure side,
the
throttle
valve
closes
the
increased pressure br-er1-09.cdr
© ISF 2002
onto
Single Pressure Reducing Valve, Shut Down
the
brane.
Figure 1.9
The
injector-type welding torch injector or blowpipe
torch consists of a body
with
valves
and welding chamber
with
mixer tube
coupling nut mixer nozzle oxygen valve
hose connection for oxygen A6x1/4" right
welding
nozzle, Figure 1.10. injector pressure nozzle suction nozzle
By the selection of suitable
welding
chambers,
fuel gas valve
welding nozzle
the welding torch head
flame intensity can br-er1-10.cdr
be
adjusted
for
different
plate thicknesses.
Figure 1.10
torch body © ISF 2002
Welding Torch
welding
hose connection for fuel gas A9 x R3/8” left
mem-
1. Gas Welding
8
The special form of the mixing chamber guarantees highest possible safety against flashback, Figure 1.11. The high outlet speed of the escaping O2 generates a negative pressure in the acetylene gas line, in consequence C2H2 is sucked and drawn-in. C2H2 is therefore available with a very low pressure of 0.02 up to 0.05 MPa compared with O2 (0.2 up to 0.3 MPa).
acetylene oxygen acetylene
welding torch head injector nozzle coupling nut
pressure nozzle
torch body
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© ISF 2002
Injector-Area of Torch
Figure 1.11 A neutral flame adjustment allows the differentiation of three zones of a chemical reaction, Figure 1.12:
0. dark core:
escaping gas mixture
1. brightly shining centre cone:
acetylene decomposition C2H2 -> 2C+H2
2. welding zone:
1st stage of combustion 2C + H2 + O2 (cylinder) -> 2CO + H2
3. outer flame:
2nd stage of combustion 4CO + 2H2 + 3O2 (air) -> 4CO2 + 2H2O
complete reaction:
2C2H2 + 5O2 -> 4CO2 + 2H2O
1. Gas Welding
9
welding flame combustion welding nozzle centre cone welding zone 2-5
outer flame
3200°C
2500°C
1800°C
1100°C
400°C
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© ISF 2002
Figure 1.12 welding flame ratio of mixture
By changing the mixture ratio of the
excess of oxygen
normal (neutral)
excess of acetylene
volumes O2:C2H2 the weld pool can greatly be influenced, Figure 1.13. At a neutral flame adjustment the mixture ratio is O2:C2H2 = 1:1. By reason of the higher flame temperature, an excess oxygen flame might allow faster welding of steel, however, there is the risk of oxidizing (flame cutting). effects in welding of steel
Area of application: brass
foaming spattering
sparking
The excess acetylene causes the carburising of steel materials.
consequences: carburizing hardening br-er1-13.cdr
reducing
oxidizing © ISF 2002
Area of application: cast iron
Effects of the Welding Flame Depending on the Ratio of Mixture
Figure 1.13
1. Gas Welding
10 By changing the gas mixture outlet
welding flame
speed the flame can be adjusted to
balanced (neutral) flame nozzle size: for plate thickness of 2-4 mm
the heat requirements of the welding
discharging velocity and weld heat-input rate: low 2
job, for example when welding plates (thickness: 2 to 4 mm) with the welding chamber size 3: “2 to 4 mm”, Fig-
soft flame discharging velocity and weld heat-input rate: middle 3
ure 1.14. The gas mixture outlet speed is 100 to 130 m/s when using a medium or normal flame, applied to
moderate flame
at, for example, a 3 mm plate. Using a
discharging velocity and weld head-input rate: high 4
soft flame, the gas outlet speed is lower (80 to 100 m/s) for the 2 mm plate, with a hard flame it is higher (130 to 160 m/s) for the 4 mm plate.
hard flame br-er1-14.cdr
© ISF 2002
Effects of the Welding Flame Depending on the Discharge Velocity
Figure 1.14 Depending on the plate thickness are the working methods “leftward weld-
Leftward welding is applied to a plate thickness of up to 3 mm. The weld-rod dips into the molten pool from time to time, but remains calm otherwise. The torch swings a little. Advantages: easy to handle on thin plates
ing” and “rightward welding” applied, Figure 1.15. A decisive factor for the designation of the working method is the sequence of flame and welding rod as well as the manipulation of flame and welding rod. The welding direction itself is of no importance. In leftward
welding-rod
flame
welding bead
Rightward welding ist applied to a plate thickness of 3mm upwards. The wire circles, the torch remains calm. Advantages: - the molten pool and the weld keyhole are easy to observe - good root fusion - the bath and the melting weld-rod are permanently protected from the air - narrow welding seam - low gas consumption
welding the flame is pointed at the open gap and “wets” the molten pool; the heat input to the molten pool can be well controlled by a slight move-
weld-rod
flame
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ment of the torch (s = 3 mm).
© ISF 2002
Flame Welding
Figure 1.15
1. Gas Welding
11
In rightward welding the flame is directed onto the molten pool; a weld
can be applied to a plate thickness of
1,5
approx. 1.5 mm without filler material,
symbol
flange weld
1,0
but this does not apply to any other
plain butt weld
1,0
4,0
3,0
12,0
1,0
8,0
1,0
8,0
lap seam
1,0
8,0
fillet weld
plate thickness and weld shape, Figure 1.16.
denotation
s
gap preparations
r=
Flanged welds and plain butt welds
plate thickness range s [mm] from to
~ ~ s+1
keyhole is formed (s = 3 mm).
V - weld 1-2 1-2
corner weld
By the specific heat input of the different welding methods all welding positions can be carried out using the oxyacetylene welding method, Figures 1.17 and 1.18
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© ISF 2002
Gap Shapes for Gas Welding
Figure 1.16 butt-welded seams in gravity position
When working in tanks and confined
PA
spaces, the welder (and all other per-
gravity fillet welds
sons present!) have to be protected against the welding heat, the gases
PB
produced during welding and lack of
horizontal fillet welds vertical fillet and butt welds
oxygen ((1.5 % (vol.) O2 per 2 % (vol.)
s
f
C2H2 are taken out from the ambient atmosphere)), Figure 1.19. The addi-
PF PG
vertical-upwelding position vertical-down position
PC
horizontal on vertical wall
PE
overhead position
PD
horizontal overhead position
tion of pure oxygen is unsuitable (explosion hazard!).
© ISF 2002
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Welding Positions I
Figure 1.17
1. Gas Welding
12 A special type of autogene method is flame-straightening, where specific locally applied flame heating allows for shape correction of workpieces, Figure
PA
1.20. Much experience is needed to
PB PF
carry out flame straightening processes. The basic principle of flame straightening depends on locally applied heating in
PC
connection with prevention of expansion. This process causes the appearance of a PG PD
heated zone. During cooling, shrinking forces are generated in the heated zone
PE
and lead to the desired shape correction. © ISF 2002
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Welding Positions II
Figure 1.18
Safety in welding and cutting inside of tanks and narrow rooms
Flame straightening
welded parts
first warm up both lateral plates, then belt
Hazards through gas, fumes, explosive mixtures, electric current protective measures / safety precautions 1. requirement for a permission to enter 2. extraction unit, ventilation
butt weld 3 to 5 heat sources close to the weld-seam
3. second person for safety reasons 4. illumination and electric machines: max 42volt
double fillet weld 1,3 or 5 heat sources
5. after welding: Removing the equipment from the tank
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© ISF 2002
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Gas Welding in Tanks and Narrow Rooms
Figure 1.19
© ISF 2002
Flame Straightening
Figure 1.20
2. Manual Metal Arc Welding
2003
2. Manual Metal Arc Welding
13 Figure 2.1 describes the burn-off of a covered stick electrode. The stick electrode consists of a core wire with a mineral covering. The welding arc between the electrode and the workpiece melts core wire and covering. Droplets of the liquefied core wire mix with the molten base material forming weld metal while the molten covering is forming slag which, due to its lower density, solidifies on the weld pool. The slag layer and gases which are generated inside the arc protect the metal during transfer and also the c ISF 2002
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Weld Point
weld pool from the detrimental influences of the surrounding atmosphere.
Figure 2.1
Covered stick electrodes
have
1. Conductivity of the arc plasma is improved by
re-
a) ease of ignition b) increase of arc stability
placed the initially
2. Constitution of slag, to
applied metal arc and
carbon
electrodes.
a) influence the transferred metal droplet b) shield the droplet and the weld pool against atmosphere c) form weld bead
arc The
3. Constitution of gas shielding atmosphere of
covering has taken on
the
a) organic components b) carbides
functions
4. Desoxidation and alloying of the weld metal
which are described
5. Additional input of metallic particles
in Figure 2.2.
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© ISF 2002
Task of Electrode Coating
Figure 2.2
2. Manual Metal Arc Welding
14
The covering of the stick electrode consists of a multitude of components which are mainly mineral, Figure 2.3. coating raw material
effect on the welding characteristics
quartz - SiO2
to raise current-carrying capacity
rutile -TiO2
fluorspar - CaF2
to increase slag viscosity, good re-striking to refine transfer of droplets through the arc to reduce arc voltage, shielding gas emitter and slag formation to increase slag viscosity of basic electrodes, decrease ionization
calcareous- fluorspar K2O Al2O3 6SiO2
easy to ionize, to improve arc stability
ferro-manganese / ferro-silicon
deoxidant shielding gas emitter
magnetite - Fe3O4 calcareous spar -CaCO3
cellulose kaolin Al2O3 2SiO2 2H2O
lubricant
potassium water glass K2SiO3 / Na2SiO3
bonding agent
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© ISF 2002
Influence of the Coating Constituents on Welding Characteristics
Figure 2.3 For the stick electrode manufacturing mixed ground and screened covering materials are used as protection for the core wire which has been drawn to finished diameter and subsequently cut to size, Figure 2.4.
raw material storage for flux production raw wire storage jaw crusher
1
wire drawing machine and cutting system 2
3
descaling magnetic separation
inspection
example of a three-stage wire drawing machine drawing plate
cone crusher for pulverisation
Ø 6 mm
sieving to further treatment like milling, sieving, cleaning and weighing
sieving system
Ø 5,5 mm
Ø 4 mm
weighing and mixing inspection
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electrode compound
3,25 mm
wet mixer
inspection © ISF 2002
Stick Electrode Fabrication 1
Figure 2.4
to the pressing plant
2. Manual Metal Arc Welding
15
the pressing plant
inspection electrodepress
electrode compound
inspection compound
packing inspection
TO DELIVERY
core wire magazine
nozzleconveying wire wire pressing belt feeder magazine head
drying stove inspection inspection inspection © ISF 2002
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Stick Electrode Fabrication 2
Figure 2.5
The core wires are coated with the covering material which contains bind-
pressing cylinder
core rod coating pressing nozzle pressing cylinder
pressing mass
core rod guide
ing agents in electrode extrusion presses. The defect-free electrodes then pass through a drying oven and are, after a final inspection, automatically packed, Figure 2.5.
Figure 2.6 shows how the moist extruded covering is deposited onto the core wire inside an electrode extrusion press.
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Production of Stick Electrodes
Figure 2.6
2. Manual Metal Arc Welding
16
Stick electrodes are, according to their covering compositions, categorized into four different types, Figure 2.7. with concern to burn-off characteristics and achievable weld metal toughness these types show fundamental differences.
cellulosic type
acid type
cellulose 40 rutile TiO2 20 quartz SiO2 25 Fe - Mn 15 potassium water glass almost no slag droplet transfer : medium- sized droplets toughness value: good
basic typ
rutile type
magnetite Fe3O4 50 SiO2 20 quartz CaCO3 10 calcite Fe - Mn 20 potassium water glass slag solidification time: long droplet transfer : fine droplets to sprinkle toughness value:
rutile TiO2 45 magnetite Fe3O4 10 SiO2 quartz 20 CaCO3 10 calcite Fe - Mn 15 potassium water glass
fluorspar CaF2 45 CaCO3 40 calcite SiO2 10 quartz 5 Fe - Mn potassium water glass
slag solidification time: medium
slag solidification time: short
droplet transfer : medium- sized to fine droplets toughness value:
droplet transfer : medium- sized to big droplets toughness value:
good
very good
normal
© ISF 2002
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Characteristic Features of Different Coating Types
Figure 2.7
The melting characteristics of the different coverings and the slag properties result in further properties; these determine the areas of application, Figure 2.8.
coating type symbol
cellulosic type C
acid type A
rutile type R
basic type B
~/+
~/+
~/+
=/+
very good
moderate
good
good
PG,(PA,PB, PC,PE,PF)
PA,PB,PC, PE,PF,PG
PA,PB,PC, PE,PF,(PG)
PA,PB,PC, PE,PF,PG
low
high
low
very low
moderate
good
good
moderate
slag detachability
good
very good
very good
moderate
characteristic features
spatter, little slag, intensive fume formation
high burn-out losses
universal application
low burn-out losses hygroscopic predrying!!
current type/polarity gap bridging ability welding positions sensitivity of cold cracking weld appearance
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© ISF 2002
Characteristics of Different Coating Types
Figure 2.8
2. Manual Metal Arc Welding
17
The dependence on temperature of the slag’s electrical conductivity determines the reignition behaviour of a stick electrode, Figure 2.9. The electrical conductivity for a rutile stick electrode lies, also at room temperature, above the threshreignition threshold
old value which is
h ac co igh id s n d - te l a uc mp g to e r r a tu re hig bas h- ic s co tem lag nd pe uc ra to tur r e
conductivity
g slag ntainin o c le ti high ru r nducto semico
necessary for reignition.
Therefore,
rutile
electrodes
are given prefertemperature
ence
in
the
© ISF 2002
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production of tack
Conductivity of Slags
welds where reignition occurs fre-
Figure 2.9
quently. The complete designation
for
filler
materials, following
DIN EN 499 - E 46 3 1Ni B 5 4 H5 3
European dardisation, cludes
hydrogen content < 5 cm /100 g welding deposit butt weld: gravity position fillet weld: gravity position suitable for direct and alternating current recovery between 125% and 160% basic thick-coated electrode chemical composition 1,4% Mn and approx. 1% Ni o minimum impact 47 J in -30 C 2 minimum weld metal deposit yield strength: 460 N/mm distinguishing letter for manual electrode stick welding
Stanin-
details–
partly as encoded abbreviation
–
which are relevant
The mandatory part of the standard designation is: EN 499 - E 46 3 1Ni B
for welding, Figure © ISF 2002
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2.10. The identification letter
Designation Example for Stick Electrodes
for the
welding process is
Figure 2.10
first: E
-
manual electrode welding
G
-
gas metal arc welding
T
-
flux cored arc welding
W
- tungsten inert gas welding
S
-
submerged arc welding
2. Manual Metal Arc Welding
18
The identification numbers give information about yield point, tensile strength and elongation of the weld metal where the tenfold of the identification number is the minimum yield point in N/mm², Figure 2.11.
key number
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
*) L0 = 5 D0
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© ISF 2002
Characteristic Key Numbers of Yield Strength, Tensile Strength and Elongation
Figure 2.11
The identification figures for the minimum impact energy value of 47 J – a parameter for the weld metal toughness – are shown in Figure 2.12.
characteristic figure Z A 0 2 3 4 5 6 7 8
0
minimum impact energy 47 J [ C] no demands +20 0 -20 -30 -40 -50 -60 -70 -80
The minimum value of the impact energy allocated to the characteristic figures is the average value of three ISO-V-Specimen, the lowest value of whitch amounts to 32 Joule. br-er2-12.cdr
Characteristic Key Numbers for Impact Energy
Figure 2.12
2. Manual Metal Arc Welding
19 The
chemical
composition
of
the weld metal is shown by the alloy symbol,
Figure
2.13.
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© ISF 2002
Alloy Symbols for Weld Metals Minimum Yield Strength up to 500 N/mm2
Figure 2.13
The properties of a stick electrode are characterised by the covering thickkey letter
ness and the covering type. Both de-
type of coating
tails are determined by the identification letter for the electrode covering, Figure 2.14.
A
acid coating
B
basic coating
C
cellulose coating
R
rutile coated (medium thick)
RR
rutile coated (thick)
RA
rutile acid coating
RB
rutile basic coating
RC
rutile cellulose coating
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Figure 2.14
© ISF 2002
2. Manual Metal Arc Welding Figure
2.15
20
ex-
plains the additional identification figure for electrode recovery and applicable type
of
The
current.
subsequent
identification figure determines the application
possibili-
ties
different
for
© ISF 2002
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Additional Characteristic Numbers for Deposition Efficiency and Current Type
welding positions: Figure 2.15 1-
all positions
2-
all positions, except vertical down position
3-
flat position butt weld, flat position fillet weld, horizontal-, vertical up position
4-
flat position butt and fillet weld
5-
as 3; and recommended for vertical down position
The last detail of the European Standard designation determines the maximum hydrogen content of the weld metal in cm³ per 100 g weld metal. Welding
current
amperage
and
core wire diameter
of
the
stick
electrode are determined
by
the
thickness
of
the
workpiece
to
be
welded. Fixed stick electrode br-er2-16.cdr
© ISF 2002
Size and Welding Current of Stick Electrodes
Figure 2.16
are each
lengths
assigned
to
diameter,
Figure 2.16.
2. Manual Metal Arc Welding
21
Figure 2.17 shows electrode holder
the process principle
of
manual
metal arc welding.
stick electrode
- (+)
Polarity and type of current depend on
power source = or ~
the
+ (-)
applied
elec-
arc
trode
types.
known
All
power work piece
sources with a de-
© ISF 2002
br-er2-17.cdr
scending
Principle Set-up of MMAW Process
characteristic curve can be used.
Figure 2.17
Since in manual metal arc welding the arc length cannot always be kept constant, a steeply descending power source is used. Different arc lengths
power source characteristic
lead therefore to just minimally altered A2
U
A1
weld current intensities, Figure 2.18. Penetration remains basically unaltered.
A2
2 1
A1
21 characteristic of the arc br-er2-18.cdr
Figure 2.18
I
© ISF 2002
2. Manual Metal Arc Welding
22 Simple welding transformers are arc welding converter
used for a.c. welding. For d.c. welding mainly converters, rectifiers and series regulator transistorised power sources (inverters) are applied. Con-
transformer
verters are specifically suitable for site
welding
and
are
mains-
independent when an internal combustion engine is used. The advanrectifier
tages of inverters are their small size and low weight, however, a more complicated electronic design is necinverter type
br-er2-19.cdr
essary, Figure 2.19.
© ISF 2002
Figure 2.19 45 RA73
Figure 2.20 shows the standard weld-
V
ing parameters of different stick elec-
40
trode diameters and stick electrode
The rate of deposition of a stick electrode is, besides the used current intensity, dependent on the so-called
medium weld voltage
types.
RR73
“electrode recovery”, Figure 2.21. This
35
RR12 RA12
30
B53
B15
25 = = = =
describes the mass of deposited weld metal / mass of core wire ratio 20
in percent. Electrode recovery can
100
200
300
3,25 4 5 6
A
400
medium weld current
reach values of up to 220% with metal br-er2-20.cdr
covering components in high-efficiency electrodes. Figure 2.20
© ISF 2002
2. Manual Metal Arc Welding
23
A survey of the material spectrum which is suitable for manual metal arc welding is given in Figure 2.22. The survey comprises almost all metals known for technical applications and also explains the wide application range of the method. cy
7
constructional steels shipbuilding steels high-strength constructional steels boiler and pressure vessel steels austenitic steels creep resistant steels austenitic-ferritic steels (duplex) scale resistant steels wear resistant steels hydrogen resistant steels high-speed steels cast steels combinations of materials (ferritic/ austenitic)
cast iron:
cast iron with lamella graphite cast iron with globular graphite
nickel:
pure nickel Ni-Cu-alloys Ni-Cr-Fe-alloys Ni-Cr-Mo-alloys
copper:
electrical grade copper (ETP copper) bronzes (CuSn, CuAl) gunmetal (CuSnZnPb) Cu-Ni-alloys
aluminium:
pure aluminium AlMg-alloys AlSi -alloys
eff ic ion
6
c
de po s it de po s it io n
ef fic
X
4
ien
cy
22 0%
5
3
16 0%
burn-off rate at 100% duty cycle
steel:
ien
kg/h
th
d te oa c ick
2 th
in-
b
a
ed at co
1 = RR12 - 5 mm RR73 - 5 mm
X=
0
0
100
200 300 welding amperage
400 A 500
a = A- and R- coated electrodes, recovery 105% b = basic-coated electrodes, recovery 2CO2 in the workpiece proximity) intensifies
wire elektrodes
this effect when CO2 is used. In argon, the current-carrying arc core
current-carrying arc core
is wider and envelops the wire electrode end, Figure 5.13. This generates electromagnetic forces which
argon
bring about the detachment of the
carbon dioxide
liquid electrode material. This socalled “pinch effect” causes a metal transfer in small drops, Figure 5.14.
The pointed shape of the arc attachbr-er5-14e.cdr
© ISF 2002
ment in carbon dioxide produces a reverse-direction
force
component,
i.e., the molten metal is pushed up Figure 5.14
until gravity has overcome that force component and material transfer in the form of very coarse drops appear.
acceleration due to gravity wire electrode
electromagnetic force FL (pinch effect)
Besides the pinch effect, the inertia and
the
gravitational
force,
other
forces, shown in Figure 5.15, are ac-
viscosity surface tension S
droplets necking down
tive inside the arc space; however these forces are of less importance.
backlash forces fr of the evaporating material
inertia electrostatic forces
suction forces, plasma flow induced work piece br-er5-15e.cdr
© ISF 2002
Forces in Arc Space
Figure 5.15
5. Gas-Shielded Metal Arc Welding
63
If the welding voltage and the wire feed speed are further increased, a rotating arc occurs after an undefined transition zone, Figure 5.16. High-efficiency MAG welding has been applied since the beginning of the nineties; the deposition rate, when this process is used, is twice the size as, in comparison, to spray arc welding. Apart from a multicomponent gas with a helium
proportion,
also a high-rating power source and a precisely controlled wire feed system for high
wire
feed
speeds are necessary.
br-er5-16e.cdr
© ISF 2002
Rotating Arc
Figure 5.16
Figure 5.17 depicts the deposition rates over the wire feed speed, as achievable with modern high-efficiency MAG welding processes.
During Ø 1,2 mm
kg/h
deposition rate
transi-
tion from the short
25
to the spray arc the
high performance GMA welding
20
Ø 1,0 mm
15
drop frequency rate increases erratically
10
Ø 0,8 mm
conventional GMA
while the drop volume
5
the 0
the
0
5
10
15
20
25
30
35
40
45 m/min
wire feed speed br-er5-17e.cdr
decreases same
degree.
With an increasing CO2-content,
this
© ISF 2002
“critical Deposition Rate
Figure 5.17
at
current
range” moves up to higher power ranges
5. Gas-Shielded Metal Arc Welding
64
and is, with inert gas constituents of lower than 80%, hardly achievable thereafter. This effect facilitates the pulsed-arc welding technique, Figure 5.18. 300
300
200
100
100
V arc voltage
200 critical current range
UEff
3
10 cm
drop volume
number of droplets
35 -4
1/s
25 20 Um
15 10 5
0
500
0 0
A
400
tP
200
600
A 400 welding current
Ikrit
Im
- background current IG - pulse voltage UP - impulse time tP - background time tG or frequency f with f = 1 / ( tG + tP), resp. - wire feed speed vD
time
IG
tG
Setting parameters:
350 300 IEff
250 200
Im
150 100 50 0 5
0
br-er5-18e.cdr
© ISF 2002
10
15 time
20
ms
br-er5-19e.cdr
30 © ISF 2002
Pulsed Arc
Figure 5.18
Figure 5.19
In pulsed-arc welding, a change-over occurs between a low, subcritical background current and a high, supercritical pulsed current. During the background phase which
welding current
corresponds with the pulsed current intensity
short arc range, the
Non-short-circuiting metal tranfer range
arc length is ionised and
backround current intensity
wire
electrode
and work surface are preheated. During the time
pulsed material
phase is
the
molten
and, as in spray arc welding,
superseded
© isf 2002
br-er5-20e.cdr
by
the
magnetic
Pulsed Metal Transfer
forces. Figure 5.20. Figure 5.20
5. Gas-Shielded Metal Arc Welding
65
Figure 5.19 shows an example of pulsed arc real current path and voltage time curve. The formula for mean current is:
Im =
1T idt T ∫0
for energy per unit length of weld is:
1T 2 i dt T ∫0
Ieff =
By a sensible se-
50 working range welding current / arc voltage
lection of welding
45
parameters,
40 optimal setting lower limit upper limit
35 voltage [v]
spray arc
GMA
the
welding
technique allows a
30 transition arc
selection of differ-
25 short arc shielding gas: 82%Ar, 18%CO2 wire diameter: 1,2 mm wire type: SG 2
20 15 10 50
75
100
125
150
175 200 225 250 welding current
275
300
325
350
375
are
distinguished
by
their
metal
400
transfer way. Fig-
© ISF 2002
br-er5-21e.cdr
ent arc types which
ure 5.21 shows the
Parameter Setting Range in GMA Welding
setting range for a
Figure 5.21
good
welding
process in the field filler metal: SG2 -1,2 mm shielding gas: Ar/He/CO2/O2-65/26,5/8/0,5
conventional
GMA welding.
transition zones spray arc
V
voltage
of
rotating arc
50
Figure 5.22 shows
30 high-efficiency spray arc
the extended set-
20
ting range for the
high-efficiency short arc
10
short arc
high-efficiency MAGM
100 br-er5-22e.cdr
200
300 welding current
400
A Quelle: Linde, ISF2002
Setting Range or Welding Parameters in Dependence on Arc Type
Figure 5.22
welding
600
process
with
rotating arc.
a
5. Gas-Shielded Metal Arc Welding
66 Some typical ap-
arc types welding methods MAGC MAGM MIG seam type, positions workpiece thickness
applications
spray arc
short arc
long arc
-
aluminium copper steel unalloyed, lowalloy, high-alloy
fillet welds or inner passes and cover passes of butt welds at medium-thick or thick components in position PA, PB
aluminium copper
different arc types
steel unalloyed, low-alloy
steel unalloyed, low-alloy, steel low-alloy, high-alloy high-alloy
steel unalloyed, low-alloy
steel unalloyed, low-alloy
fillet welds or inner passes and cover passes of butt welds at medium-thick or thick components in position PA, PB
fillet welds or butt welds fillet welds or inner at thin sheets, all positions passes and cover passes of thin and root layers of butt welds medium-thick at medium-thick or thick components, all components, all positions positions
welding of root layers in position PA
plications of the
pulsed arc
aluminium (s < 1,5 mm)
are depicted in Fig-
-
inner passes and cover passes of fillet or butt welds in position PC, PD, PE, PF, PG (out-of-position)
ure
5.23.
The
rotating arc, (not mentioned in the figure), is applied
root layer welds only conditionally possible
in just the same way as the spray
br-er5-23e.cdr
© ISF 2002
arc, however, it is
Applications of Different Arc Types
not used for the Figure 5.23
welding of copper and aluminium.
The arc length within the working range is linearly dependent on the set
U
welding voltage, Figure 5.24. The AL
weld seam shape is considerably in-
AM
AK
arc length: long medium short
fluenced by the arc length. A long arc produces a wide flat weld seam and, in the case of fillet welds, generally undercuts. A short arc produces a narrow, banked weld bead.
On the other hand, the arc length is inversely proportional to the wire
vD, I
operating point: wire feed speed: arc length: welding current: deposition efficiency:
AL
AM
AK
low long low low
medium medium medium medium
high short high high
weld appearance:
feed speed, Figure 5.25. This has influence on the current over the internal adjustment with a slightly dropping power
source
characteristic.
br-er5-24e.cdr
This
Wire Feed Speed
again is of considerable importance for the deposition rate, i.e., a low wire feed speed leads to a low deposition
© ISF 2002
Figure 5.24
5. Gas-Shielded Metal Arc Welding
67 rate, the result is flat penetration and
arc length: long medium short
U AL AM AK
low base metal fusion. At a constant weld speed and a high wire feed speed a deep penetration can be obtained.
vD, I
operating point: welding voltage: arc length:
AL
AM
high long
medium medium
AK low short
At equal arc lengths, the current intensity is dependent on the contact tube distance, Figure 5.26. With a large contact tube distance, the wire
weld appearance butt weld
stickout is longer and is therefore characterised by a higher ohmic resisweld appearance fillet weld
tance which leads to a decreased current intensity. For the adjustment of
br-er5-25e.cdr
© ISF 2002
Welding Voltage
the contact tube distance, as a thumb rule, ten to twelve times the size of
Figure 5.25 the wire diameter should be considered. lk1
lk2
lk3
influence on weld formation and welding process, Figure 5.27. When welding with the torch pointed in forward direction of the weld, a part of the weld pool is moved in front of the arc. This results in process instability.
contact tube-to-work distance lk
The torch position has considerable 3
30 mm
2
20
lk = 10 to 12 dD 1
10
0 200
250
However, it ha s the advantage of a flat smooth weld surface with good gap bridging. When welding with the torch pointed in reversing direction of
operating rule:
300 A
350
current wire electrode:
1,2 mm diameter
shielding gas:
82% Ar + 18% CO2
arc voltage:
29 V
wire feed speed:
8,8 m/min
welding speed:
58 cm/min
br-er5-26e.cdr
the weld, the weld process is more
© ISF 2002
Contact Tube-to-Work Distance
stable and the penetration deeper, as Figure 5.26
5. Gas-Shielded Metal Arc Welding
68 base metal fusion by the arc is better,
advance direction
although the weld bead surface is irregular and banked.
Figure 5.28 shows a selection of different application areas for the GMA technique and the appropriate shieldpenetration:
shallow
average
deep
gap bridging:
good
average
bad
arc stability:
bad
average
good
spatter formation: strong
average
low
weld width:
average
narrow
average
rippled
ing gases.
The welding current may be produced by different welding power sources. In d.c. welding the transformer must be wide
equipped with downstream rectifier weld appearance: smooth
br-er5-27e.cdr
assemblies, Figure 5.29. An additional
© ISF 2002
ripple-filter choke suppresses the residual ripple of the rectified current
Torch Position
and has also a process-stabilising Figure 5.27
effect.
power
sources
became
possible,
Figure
92% Ar + 8% CO2 forming gas (N2-H2-mixture)
88% Ar + 12% O2 82% Ar + 18% CO2
application examples autoclaves, vessels, mixers, cylinders panelling, window frames, gates, grids stainless steel pipes, flanges, bends spherical holders, bridges, vehicles, dump bodies reactors, fuel rods, control devices rocket, launch platforms, satellites valves, sliders, control systems stator packages, transformer boxes passenger cars, trucks radiators, shock absorbers, exhausts cranes, conveyor roads, excavators (crawlers) shelves (chains), switch boxes braces, railings, stock boxes mud guards, side parts, tops, engine bonnets attachments to flame nozzles, blast pipes, rollers vessels, tanks, containers, pipe lines stanchions, stands, frames, cages beams, bracings, craneways harvester-threshers, tractors, narrows, ploughs waggons, locomotives, lorries
5.29. The operating principle of a transistor
80% Ar + 5% O2 + 15% CO2 92% Ar + 8% O2
industrial sections
analogue
83% Ar + 15% He + 2% CO2 90% Ar + 5% O2 + 5% CO2
sign of transistor
99% Ar + 1% O2 or 97% Ar + 3% O2 97,5% Ar + 2,5% CO2
transistors the de-
Argon 4.8 Helium 4.6
efficient Argon 4.6
of
shielding gases
ment
Ar/He-mixture Ar + 5% H2 or 7,5% H2
With the develop-
analogue br-er5-28e.cdr
power source fol-
Fields of Application of Different Shielding Gases
lows the principle of an audio frequency
© ISF 2002
Figure 5.28
amplifier which amplifies a low-level to a high level input signal, possibly distortion-free. The transistor power source is, as conventional power sources, also equipped with a three-phase
5. Gas-Shielded Metal Arc Welding
69
transformer, with generally only one secondary tap. The secondary voltage is rectified by silicon diodes into full wave operation, smoothed by capacitors and fed to the arc through a transistor cascade. The welding voltage is steplessly adjustable until no-load voltage is reached. The difference between source voltage and welding voltage reduces at the transistor cascade and produces a comparatively high stray power which, in general, makes water-cooling necessary. The efficiency factor is between 50 and 75%. This disadvantage is, however, accepted as those power sources are characterised by very short reaction times (30 to 50 µs). Along with the development of transistor analogue power sources, the consequent separation of the power section (transformer and rectifier) and electronic control took place. The analogue or digital control sets the reference values and also controls the welding process. The power section operates exclusively as an amplifier for the signals coming from the control.
The output stage may also be carried out by clocked cycle. A secondary clocked transistor power source features just as the analogue power sources, a transformer and a rectifier, Figure 5.30. The transistor unit functions as an on-off switch. By varying the on-off period, i.e., of the pulse duty factor, the average voltage at the output of the transistor stage may be varied. The arc voltage achieves small ripples, which are of a limited amplitude, in the switching frequency of, in general, 20 kHz; whereas the welding current shows to be strongly smoothed during the high pulse frequencies caused by inductivities. As the transistor unit has only a switching function, the stray power is lower than that three-phase transformer
fully-controlled three-phase bridge rectifier
energy store
of
analogue
sources. The effi-
transistor power section
mains supply
welding current
ciency factor is approx. 75 – 95%. The reaction times of
uist u1 . . un
reference input values
iist
signal processor (analog-to-digital)
these
clocked
units are within of current pickup
300
–
500
µs
clearly longer than © isf 2002
br-er5-29e.cdr
GMA Welding Power Source, Electronically Controlled, Analogue
Figure 5.29
those of analogue power sources.
5. Gas-Shielded Metal Arc Welding
70
Series regulator power sources, the so-called “inverter power sources”, differ widely from the afore-mentioned welding machines, Figure 5.31. The alternating voltage coming from the mains (50 Hz) is initially rectified, smoothed and converted into a medium frequency alternating voltage (approx. 25-50 kHz) with the help of controllable transistor and thyristor switches. The alternating voltage is then transformer reduced to welding voltage levels and fed into the welding process through a secondary rectifier, where the alternating voltage also shows switching frequency related ripples. The advantage of inverter power sources is their low weight. A transformer that
transforms
voltage
with
fre-
quency of 20 kHz, has, compared with a
50
former,
Hz
3-phase transformer
3-phase bridge rectifier
energy store
transistor switch
protective reactor welding current
mains supply
trans-
considera-
bly lower magnetic
Uist U1 . . Un
losses, that is to
reference input values
say, its size may accordingly
be
smaller
its
Iist
signal processor (analog-to-digital)
br-er5-30e.cdr
and
© ISF 2002
GMA Welding Power Source, Electronically Controlled, Secondary Chopped
weight is just 10% of that of a 50 Hz
current pickup
Figure 5.30
transformer.
Reaction time and efficiency
factor
are comparable to the
filter
3-phase bridge rectifier
energy storage
transistor inverter
medium frequency transformer
rectifier welding current
mains supply
corresponding
values of switchingUist
type power sources.
U1 . . Un
reference input values
br-er5-31e.cdr
Iist
signal processor (analog-to-digital)
current pickup
© ISF 2002
GMA Welding Power Source, Electronically Controlled, Primary Chopped, Inverter
Figure 5.31
5. Gas-Shielded Metal Arc Welding
71
All welding power sources are fitted with a rating plate, Figure 5.32. Here the performance capability and the properties of the power source are listed. The S in capital letter (former K) in manufacturer insulations class
rotary current welding rectifier
~
_
protective IP21 system
VDE 0542 production number
type welding MIG/MAG
U0 15 - 38 V
F
cooling type
the middle shows F
that
DIN 40 050
input 3~50Hz 6,6 kVA (DB) cosj 0,72
power
source is suitable
switchgear number
S
the
35A/13V - 220A/25V
power range
X 60% ED 100% ED 170 A I2 220 A 23 V U2 25 V
power capacity in dependence of current flow
17 A 10 A
U1 220 V
I1 26 A
U1 380 V
I1
15 A
U1
V
I1
A
A
U1
V
I1
A
A
power supply
for welding operations
under
ardous
haz-
situations,
i.e., the secondary no-load voltage is lower than 48 Volt
min. and max. no-load voltage © ISF 2002
br-er5-32e.cdr
and therefore not Rating Plate
dangerous to the welder.
Figure 5.32
Besides the familiar solid wires also filler wires are used for
gas-shielded
metal arc welding. They consist of a a
seamless flux-cored wire electrode
b
c
metallic tube and a flux
form-enclosed flux-cored wire electrode
core
Figure 5.33 depicts common
br-er5-33e.cdr
cross-
© ISF 2002
Cross-Sections of Flux-Cored Wire Electrodes
Figure 5.33
filling.
sectional shapes.
5. Gas-Shielded Metal Arc Welding
72
Filler wires contain arc stabilisators, slag-forming and also alloying elements which support a stable welding process, help to protect the solidifying weld from the atmosphere and, more often than not, guarantee very good mechanical properties. An important distinctive criteria is the type of the filling. The influence of the filling is symbol R
slag characteristics rutile base, slowly soldifying slag rutile base, rapidly soldifying slag basic filling: metal powder
P B M V W
rutile- or fluoride-basic fluoride basic, slowly soldifying slag fluoride basic, slowly soldifying slag other types
Y S
customary application* S and M
very similar to that shielding gas ** C and M2
S and M
C and M2
S and M S and M S S and M
C and M2 C and M2 without without
S and M
without
Figure 5.34
electrode
covering in manual electrode (see
welding
chapter
2).
Figure 5.34 shows a list of the differ-
wire. © ISF 2002
Type Symbols of Flux-Cored Wire Electrodes According to DIN EN 12535
the
ent types of filler
*) S: single pass welding - M: multi pass welding **) C: CO2 - M2: mixed gas M2 according to DIN EN 439 br-er5-34e.cdr
of
6. Narrow Gap Welding, Electrogas - and Electroslag Welding
2003
6. Narrow Gap Welding, Electrogas- and Electroslag Welding
73
Up to this day, there is no universal agreement about the definition of the term “Narrow Gap Welding” although the term is actually self-explanatory. In the international technical literature, the process characteristics mentioned in the upper part of Figure 6.1 are frequently connected with the definition for narrow gap welding. In spite of these Process characteristics: - narrow, almost parallel weld edges. The small preparation angle has the function to compensate the distortion of the joining members - multipass technique where the weld build-up is a constant 1 or 2 beads per pass - usually very small heat affected zone (HAZ) caused by low energy input
“definition
difficulties” questions the
Disadvantages - higher apparatus expenditure, espacially for the control of the weld head and the wire feed device - increased risk of imperfections at large wall thicknesses due to more difficult accessibility during process control - repair possibilities more difficult
about
universally
valid Advantages: - profitable through low consumption quantities of filler material, gas and/ or powder due to the narrow gaps - excellent quality values of the weld metal and the HAZ due to low heat input - decreased tendency to shrink
all
advantages
and disadvantages of the narrow gap welding
method
can be clearly answered.
© ISF 2002
br-er6-01e.cdr
Narrow Gap Welding
Figure 6.1 The numerous variations of narrow gap welding are, in general, a further development of the conventional welding technologies. Figure 6.2 shows a classification with emphasis on several important process characteristics. Narrow gap TIG welding with cold or hot wire addition is mainly applied as an orbital process method or for the joining submerged arc electroslag narrow narrow gap welding gap welding process with straightened wire electrode (1R/L, 2R/L, 3R/L) process with oscillating wire electrode (1R/L) process with twin electrode (1R/L, 2R/L) process with lengthwise positioned strip electrode (2R/L) flat position
gas-shielded metal arc narrow gap welding
tungsten innert gas-shielded narrow gap welding
high-
alloy as well as non-ferrous
process with linearly oscillating filler wire
process with stripshaped filler and fusing feed
electrogas process with linearly oscillating wire electrode electrogas process with bent, longitudinally positioned strip electrode
process with hot wire addition (1R/L, 2R/L) MIG/MAGprocesses (1R/L,2R/L,3R/L)
als. This method is, however, hardly
tice.
The
other
processes more
vertical up position
met-
applied in the pracprocess with cold wire addition (1R/L, 2R/L)
all welding positions
br-er6-02e.cdr
spread
are widely
and
explainedin Survey of Narrow Gap Welding Techniques Based on Conventional Technologies
Figure 6.2
of
are detail
in the following.
6. Narrow Gap Welding, Electrogas- and Electroslag Welding
74
In Figure 6.3, a systematic subdivision GMA narrow gap welding no wire-deformation
GMA narrow gap welding wire-deformation
technologies is shown. In accor-
long-wire method (1 B/P, 2 B/P) thick-wire method (1 B/P, 2 B/P) twin-wire method (1 B/P)
D
A
whether the process is carried out B
coiled-wire method (1 B/P) corrugated wire method with mechanical oscillator (1 B/P) corrugated wire method with oscillating rollers (1 B/P) corrugated wire method with contour roll (1 R/L) zigzag wire method (1 B/P) wire loop method (1 B/P)
explanation: B/ P: Bead/ Pass
A: method without forced arc movement B: method with rotating arc movement C: method with oscillating arc movement D: method with two or more filler wires
br-er6-03e.cdr
dance with this, the fundamental distinguishing feature of the methods is
tandem-wire method (1 B/P, 2 B/P, 3 B/P) twisted wire method (1 B/P)
rotation method (1 B/P)
of the various GMA narrow gap
with or without wire deformation. Overlaps in the structure result from the application of methods where a single or several additional wires are
C
used. While most methods are suitable for single layer per pass welding, other methods require a weld build-up with at least two layers per pass. A
© ISF 2002
further subdivision is made in accordance with the different types of arc movement.
Figure 6.3 In the following, some of the GMA narrow gap technologies are explained: Using the turning tube method, Figure 6.4, side wall fusion is achieved by the turning of the contact tube; the contact tip angles are set in degrees of between 3° and 15° towards the torch axis. With an electronic stepper motor control, arbitrary transversearc oscillating mocorrugated wire method with mech. oscillator
tions with defined
1
1
dwell periods of os-
2
2 3
cillation and oscillation frequencies can be realised - inde-
3
4
4
5
5
6
6
contrast, when the corrugated
wire
method
me-
1 - wire reel 2 - drive rollers 3 - wire mechanism for wire guidance 4 - inert gas shroud 5 - wire guide tube and shielding gas tube 6 - contact tip
1 - wire reel 2 - mechanical oscillator for wire deformation 3 - drive rollers 4 - inert gas shroud 5 - wire feed nozzle and shielding gas tube 6 - contact tip
br-er 6-04e.cdr
with
Principle of GMA Narrow Gap Welding
chanical oscillator is Figure 6.4
8 - 10
wire properties. In
12 - 14
pendent of the filler
6. Narrow Gap Welding, Electrogas- and Electroslag Welding
75
applied, arc oscillation is produced by the plastic, wavy deformation of the plate thickness: gap preparation:
300 mm square-butt joint, 9 mm flame cut 1.2 mm elctrode diameter: amperage: 260 A pulse frequency: 120 HZ arc voltage: 30 V welding speed: 22 cm/min -1 wire oscillation: 80 min oscillation width: 4 mm shielding gas: 80% Ar/ 20% Co2 primery gas flow: 25 l/min secondary gas flow: 50 l/min number of passes: approx. 70
wire electrode. The deformation is obtained by a continuously swinging oscillator which is fixed above the wire feed rollers. Amplitude and frequency of the wave motion can be varied over the total amplitude of oscillation and the speed of the mechanical oscillator or, also, over the wire feed speed. As the contact tube remains stationary, very narrow gaps with widths from 9 to 12 mm with plate thicknesses of up to 300 mm can be welded.
br-er6-05e.cdr
© ISF 2002
Figure 6.5 Figure 6.5 shows the macro section of a GMA narrow gap welded joint with plates (thickness: 300 mm) which has been produced by the mechanical oscillator method in approx. 70 passes. A highly regular weld build-up and an almost straight fusion line with an extremely narrow heat affected zone can be noticed. Thanks to the correct setting of the oscillation parame-
rotation method 1
spiral wire method 1
ters and the precise, centred torch manipulation
2 3
2 3 4
no
4
5
sidewall fusion de-
6
5
of
the
low
sidewall penetration depth. A further ad-
1 - wire reel 2 - drive rollers 3 - mechanism for nozzle rotation 4 - inert gas shroud 5 - shielding gas nozzle 6 - wire guiding tube
1 - wire reel 2 - wire mechanism for wire deformation 3 - drive rollers 4 - wire feed nozzle and shielding gas supply 5 - contact piece
br-er 6-06e.cdr
vantage
of
weave-bead
the
Principle of GMA Narrow Gap Welding
techFigure 6.6
9 - 12
spite
13 - 14
fects occurred, in
6. Narrow Gap Welding, Electrogas- and Electroslag Welding
76
nique is the high crystal restructuring rate in the weld metal and in the basemetal adjacent to the fusion line – an advantage that gains good toughness properties.
Two narrow-gap welding variations with a rotating arc movement are shown in Figure 6.6. When the rotation method is applied, the arc movement is produced by an eccentrically protruding wire electrode (1.2 mm) from a contact tube nozzle which is rotating with frequencies between 100 and 150 Hz. When the wave wire method is used, the 1.2 mm solid wire is being spiralwise deformed. This happens before it enters the rotating 3 roll wire feed device. With a turning speed of 120 to 150 revs per minute the welding wire is deformed. The effect of this is such that after leaving the contact piece the deformed wire creates a spiral diameter of 2.5 to 3.0 mm in the gap – adequate enough to weld plates with thicknesses of up to 200 mm at gap widths between 9 and 12 mm with a good sidewall fusion.
Figure 6.7 explains two GMA narrow gap welding methods which are operated without forced arc movement, where a reliable sidewall fusion is obtained either by the wire deflection through constant deformation (tandem wire method) or by forced wire deflection with the contact tip (twin-wire method). In both cases, two wire electrodes with thicknesses between 0.8 and 1.2 mm are used. When the tandem technique is applied, these electrodes are transported to the two weld heads which are arranged inside the gap in tandem and which are indeFigure pendently selectable.
When tandem method
twin-wire method
1
1
350
2
4
3
5
4
6
twin-
wire method is applied, two parallel
2 3
the
switched
elec-
trodes are transported by a com-
5
9 - 12
1 - wire reel 2 - deflection rollers 3 - drive rollers 4 - inert gas shroud 5 - shielding gas nozzle 6 - wire feed nozzle and contact tip
1 - wire reel 2 - drive rollers 3 - inert gas shroud 4 - wire feed nozzle and shielding gas supply 5 - contact tips
15 - 18
mon wire feed unit, and, subsequently, adjusted
in
a
common
sword-
br-er 6-07e.cdr
Principle of GMA Narrow Gap Welding
Figure 6.7
type torch at an incline towards the
6. Narrow Gap Welding, Electrogas- and Electroslag Welding
77
weld edges at small spaces behind
strip electrode
each other (approx. 8 mm) and molSO stick out s
α
s a x α
gap width electrode deflection distance of strip tip to flank twisting angle
h w
bead hight bead width
so x
a
h
f
ten.
In place of the SA narrow gap welding methods, mentioned in Figure 6.2, the method with a lengthwise po-
w
twin-wire electrode
sitioned strip electrode as well as the twin-wire method are explained in
vw
s
H z
a h
vw a H z
weld speed electrode deflection stick out distance torch - flank
s h w p
gap width bead height bead width penetration depth
more detail, Figure 6.8. SA narrow gap welding with strip electrode is carried out in the multipass layer technique, where the strip electrode is deflected at an angle of approx. 5°
p
w
towards the edge in order to avoid
br-er6-08e.cdr
Submerged Arc Narrow Gap Welding
collisions. After completing the first
Figure 6.8 10°
7°
fillet weld and slag removal the oppo8 s
s
8
site fillet is made. Solid wire as well as cored-strip electrodes with widths be-
16
tween 10 and 25 mm are used. The gap width is, depending on the number of passes per layer, between 20 and
double-U butt weld SA-DU weld preparation (8UP DIN 8551) 8°
square-edge butt weld SA-SE weld preparation (3UP DIN 8551) 10
25 mm. SA twin-wire welding is, in general, carried out using two elecs
3
s
6
trodes (1.2 to 1.6 mm) where one electrode is deflected towards one weld
of the groove or towards the opposite weld edge. Either a single pass layer
3
edge and the other towards the bottom double-U butt weld GMA-DU weld preparation (Indexno. 2.7.7 DIN EN 29692)
narrow gap weld GMA-NG weld preparation (not standardised)
br-er6-09e.cdr
Comparison of the Weld Groove Shape
or a two pass layer technique are applied. Dependent on the electrode diFigure 6.9
6. Narrow Gap Welding, Electrogas- and Electroslag Welding
78
ameter and also on the type of welding powder, is the gap width between 12 and 13 mm.
Figure 6.9 shows a comparison of groove shapes in relation to plate thickness for SA welding (DIN 8551 part 4) with those for GMA welding (EN 29692) and the unstandardised, mainly used, narrow gap welding. Depending on the plate thickness, significant differences in the weld crosssectional dimensions occur which may lead to substantial saving of mabr-er6-10e_sw.cdr
© ISF 2002
terial and energy during welding. For example, when welding thicknesses of 120 mm to 200 mm with the narrow
Figure 6.10 gap welding technique, 66% up to
workpiece
wire guide
75% of the weld metal weight are electrode
edge weld.
shielding gas +
saved, compared to the SA square
arc
The practical application of SA narrow
weld pool Cu-shoe weld advance
gap welding for the production of a
weld metal
water
flanged calotte joint for a reactor pressure vessel cover is depicted in Figure 6.10. The inner diameter of the pressure
vessel
is
more
than
5,000 mm with wall thicknesses of 400 mm
and
40,000 mm.
with
The
a
height
of
total
weight
is
designation: gas-shielded metal arc welding (GMAW acc. DIN 1910 T.4) position: vertical (width deviations of up to 45°) plate thickness: 6 - 30 mm square-butt joint or V weld seam 30 mm double-V weld seam materials: unalloyed, lowalloy and highalloy steels gap width: 8 - 20 mm electrodes: only 1 (flux-cored wire, for slag formation between copper shoe and weld surface) Ø 1.6 - 3.2 mm amperage: 350 - 650 A voltage: 28 - 45 V weld speed: 2 - 12 m/h shielding gas: unalloyed and lowalloy steels CO2 or mixed gas (Ar 60% and 40% Co2 ) highalloy steels: argon or helium br-er6-11e.cdr
Electrogas Welding
3,000 tons. The weld depth at the joint was 670 mm, so it had been necesFigure 6.11
6. Narrow Gap Welding, Electrogas- and Electroslag Welding
79
sary to select a gap width of at least 35 mm and to work in the three pass layer technique.
Electrogas welding (EG) is characterised by a vertical groove which is bound by two water-cooled copper shoes. In the groove, a filler wire electrode which is fed through a copper nozzle, is melted by a shielded arc, Figure 6.11. During this process, are groove edges fused. In relation with the ascending rate of the weld pool volume, the welding nozzle and the copper shoes are pulled upwards. In order to avoid poor fusion at the beginning of the welding, the process has to be started on a run-up plate which closes the bottom end of the groove. The shrinkholes forming at the weld end are transferred onto the run-off plate. If possible, any interruptions of the welding process should be avoided. Suitable power sources are rectifiers with a slightly dropping static characteristic. The electrode has a positive polarity.
The application of electrogas welding for low-alloyed steels is – more often than not limited, as the toughness of the heat affected zone with the complex coarse grain formation does not meet sophisticated demands. Long-time exposure to temperatures of more than 1500°C and
1 2 3 4 5 6 7 8 9
low crystallisation rates are responsible for this. The same applies to the weld metal. For a more wide-spread
1. base metal 2. welding boom 3. filler metal 4. slag pool 5. metal pool
application of electrogas welding, the High-Speed
Electrogas
6. copper shoe 7. water cooling 8. weld seam 9. Run-up plate
Welding
Method has been developed in the ISF. In this process, the gap crosssection is reduced and additional metal powder is added to increase the deposition rate. By the increase of the welding speed, the dwell times of weld-adjacent regions above critical
designation: position: plate thickness: gap width: materials: electrodes:
resistance fusion welding vertical (and deviation of up to 45°) 30 mm (up to 2,000 mm) 24 - 28 mm unalloyed, lowalloy and highalloy steels 1 or more solid or cored wires Ø 2.0 - 3.2 mm plate thickness range per electrode: fixed 30 - 50 mm oscillated: up to 150 mm amperage: 550 - 800 A per electrode voltage: 35 - 52 V welding speed: 0.5 - 2 m/h slag hight: 30 - 50 mm br-er6-12e.cdr
temperatures and thus the brittleness
Electroslag Welding
effects are significantly reduced. Figure 6.12
6. Narrow Gap Welding, Electrogas- and Electroslag Welding
80
Figure 6.12 shows the process principle of Electroslag Welding. Heating and melting of the groove faces occurs in a slag bath. The temperature of the slag bath must always exceed the melting temperature of the metal. The Joule effect, produced when the current is transferred through the conducting bath, keeps the slag bath temperature constant. The welding current is fed over one or more endless wire electrodes which melt in the highly heated slag bath. Molten pool and slag bath which both form the weld pool are, sideways retained by the groove faces and, in general, by water-cooled copper shoes which are, with the complete welding unit, and in relation with the welding speed, moved progressively upwards. To avoid the inevitable welding defects at the
beginning
of
~
the welding procpowder
ess
slag
penetration, incluignition with arc
powder fusion
sion of unmolten powder) and at the
slag
end of the welding
molten pool weld metal
start of welding
(insufficient
(shrinkholes, welding
end of welding
slag
inclusions), run-up
© ISF 2002
br-er6-13e.cdr
Process Phases During ES Welding
and run-off plates are used.
Figure 6.13 The electroslag welding process can be divided into four process phases, Figure 6.13. At the beginning of the welding process, in the so-called “ignition phase”, the arc is ignited for a short period and liquefies the non-conductive welding flux powder into conductive slag. The arc is extinguished as the electrical conductivity of the arc length exceeds that of the conductive slag. When the desired slag bath level is reached, the lower ignition parameters (current and voltage) are, during the so-called “Data-Increase-Phase”, raised to the values of the stationary welding process. This occurs on the run-up plate. The subsequent actual welding process starts, the process phase. At the end of the weld, the switch-off phase is initiated in the run-off plate. The solidifying slag bath is located on the run-off plate which is subsequently removed.
6. Narrow Gap Welding, Electrogas- and Electroslag Welding
81
The electroslag welding with consumable feed wire (channel-slot welding) proved to be very useful for shorter welds.
The copper sliding shoes are replaced by fixed Cu cooling bars and the welding nozzle by a steel tube, Figure 6.14. The length of the sheathed steel tube, in general, corresponds with the weld seam length (mainly shorter than 2.500 mm) and the steel tube melts during welding in the ascending slag bath. Dependent on the plate thickness, welding can be carried out with one single or with several wire and strip electrodes. A feature of this process variation is the handiness of the welding device and the easier welding drive motor
wire or strip electrode
Electroslag fusing nozzle process (channel welding)
welding cable run-off plate workpiece
workpiece
position: vertical plate thickness: 15 mm materials: unalloyed, lowalloy and highalloy steels
area
preparation.
Also curved seams can be welded with a bent consumable
= ~
welding consumables:
fusing feed nozzle run-up plate
workpiece cable
copper shoes workpiece
workpiece
wire electrodes: Ø 2.5 - 4 mm strip electrodes: 60 x 0.5 mm plate electrodes: 80 x60 up to 10 x 120 mm fusing feed nozzle: Ø 10 - 15 mm welding powder: slag formation with high electrical conductivity
electrode. As the groove width can be
significantly
reduced
when
comparing
copper shoes
with
br-er 6-14e.cdr
conventional proc-
Electroslag Welding with Fusing Wire Feed Nozzle
esses, and a strip shaped filler mate-
Figure 6.14
rial with a consumable technological measures post weld heat treatment
decrease of peak temperature and dwell times at high temperatures
metallurgical measures increase of purity
addition of suitable alloy and micro-alloy elements (nucleus formation)
increase of welding speed reduction of energy per unit length continuous normalisation furnace normalisation
increase of deposit rate
decrease of groove volume
application of several wire electrodes, metal powder addition
V, double-V butt joints, multi-pass technique
application of suitable base and filler metals
reduction of S-, P-, H2-, N2 and O2 - contents and other unfavourable trace elements
guide
piece is used, this welding process is rightly placed under the group of narrow gap weld-
C-content limits Mn, Si, Mo, Cr, Ni, Cu, Nb, V, Zr, Ti
ing techniques.
Likewise in elec-
br-er 6-15e.cdr
Possibilities to Improve Weld Seam Properties
Figure 6.15
trogas welding, the electroslag welding
6. Narrow Gap Welding, Electrogas- and Electroslag Welding
82
process is also characterised by a large molten pool with a – simultaneously - low heating and cooling rate. Due to the low cooling rate good degassing and thus almost porefree hardening of the slag bath is possible. Disadvantageous, however, is the formation of a coarse-grain structure. There are, however, possibilities to improve the weld properties, Figure 6.15.
To avoid postweld heat treatment the electroslag welding process with local continuous normalisation has been developed for plate thicknesses of up to approx. 60 mm, Figure 6.16. The welding temperature in the weld region drops below the Ar1temperature and is subsequently heated to the normalising temperature (>Ac3). The specially designed torches follow the copper
temperature °C 1. filler wire 2. copper shoes 3. slag pool 4. metal pool 5. water cooling 6. slag layer 7. weld seam 8. distance plate 9. postheating torch 10. side plate 11. heat treated zone
2 2000 1500 900
7 8 9
500 10 950
11
1 2 3 4 5 6 7 8 9 10
along
shoes the
weld
seam. By reason of the residual heat in the workpiece the necessary perature
temcan
be
reached in a short br-er 6-16e.cdr
ES Welding with Local Continuous Normalisation
time.
Figure 6.16
In order to circumvent an expensive postheat weld treatment which is often unrealistic for use on-site, the electroslag high-speed welding process with multilayer technique has been developed. Similar to electrogas welding, the weld cross-section is reduced and, by application of a twin-wire electrode in tandem arrangement and addition of metal powder, the weld speed is increased, as in contrast to the conventional technique. In the heat affected zones toughness values are determined which correspond with those of the unaffected base metal. The slag bath and the molten pool of the first layer are retained by a sliding shoe, Figure 6.17. The weld preparation is a double-V butt weld with a gap of approx. 15 mm, so the carried along sliding shoe seals the slag and the metal bath. Plate preparation is, as in conventional elec-
6. Narrow Gap Welding, Electrogas- and Electroslag Welding
83
troslag welding, exclusively done by flame cutting. Thus, the advantage of easier weld preparation can be main1
tained.
2 3 4
For larger plate thicknesses (70 to
9 5
100 mm), the three passes layer
6 7
technique
8
When welding the first pass with a
4
10
br-er6-17e.cdr
1 magnetic screening 2 metal powder addition 3 tandem electrode 4 water cooling 5 copper shoe (water cooled) 6 slag pool 7 molten pool 8 solidified slag 9 welding powder addition 10 weld seam
ES-welding in 2 passes with sliding shoe
© ISF 2002
has
double-V-groove
been
developed.
preparation
(root
width: 20 to 30 mm; gap width: approx. 15 mm) two sliding shoes which are adjusted to the weld groove are used. The first layer is welded using the conventional technique, with one wire electrode without metal powder addition.
Figure 6.17 12
When welding the outer passes flat
11
Cu shoes are again used, Figure 6.18. 1
Three wire electrodes, arranged in a
2
triangular formation, are used. Thus,
3 4
one electrode is positioned close to
9
the root and on the plate outer sides
5 6
two electrodes in parallel arrangement
7
are fed into the bath. The single as
8 4
well as the parallel wire electrodes are fed with different metal powder quantities which as outcome permit a welding speed 5 times higher than the 10
speed of the single layer conventional technique and also leads to strong
br-er6-18e.cdr
1 magnetic screening 2 metal powder supply 3 three-wire electrode 4 water cooling 5 copper shoe (water cooled) 6 slag pool 7 molten pool 8 solidified slag 9 welding powder supply 10 weld seam 11 first pass 12 second pass
ES-welding of the outer passes
increase of toughness in all zones of the welded joint. Figure 6.18
© ISF 2002
6. Narrow Gap Welding, Electrogas- and Electroslag Welding
84
If wall thicknesses of more than 100 mm are to be welded, several twin-wire electrodes with metal powder addition have to be used to reach deposition rates of approx. 200 kg/h. The electroslag welding process is limited by the possible crack formation in the centre of the weld metal. Reasons for this are the concentration of elements such as sulphur and phosphor in the weld centre as well as too fast a cooling of the molten pool in the proximity of the weld seam edges.
7. Pressure Welding
2003
7. Pressure Welding
85
Figure 7.1 shows a survey of the pressure welding processes for joining of metals in accordance with DIN 1910.
welding
pressure welding
In
fusion welding
gas
pressure
welding a distincgas pressure welding
resistance pressure welding
induction pressure welding
conductive pressure welding
resistance spot welding
projection welding
roll seam welding
friction welding
tion is made between
pressure butt welding
flash butt welding
open
square
and
closed
square
gas
pressure
br-er7-01e.cdr
welding, Classification of Welding Processes acc. to DIN 1910
Fig-
ure 7.2.
Figure 7.1 Both methods use efficient gas torches to bring the workpiece ends up to the welding temperature. When the welding temperature is reached, both joining members are butt-welded by the application of axial force when a flash edge forms. The long preheating time leads to the formation of a coarse-grained structure in the joining area, therefore the welds are of low toughness values. As the process is operated mainsindependently and initial state: gap closed
initial state: gap opened
the process equip-
(for special cases)
gas flame torch in the open gap stationary
ment weight
mobile
is
low
and
in also
easy to handle, the workpiece closed gap
ring-shaped burner (sectional view) pressure
1. heating
main
application
areas of gas pres-
2. torch positioning 3. welding by rapid pressing
completed weld seam working cycles: 1. heating 2. welding by pressing
sure welding are the welding of reinforcement
steels
br-er7-02e.cdr
Open Square and Closed Square Gas Pressure Welding
Figure 7.2
and of pipes in the building trade.
7. Pressure Welding
86
In pressure butt welding, the input of the necessary welding heat is produced by resistance heating. The necessary axial force is applied by copper clamping jaws which also receive the current supply, Figure 7.3. The current circuit is closed over the abutting surfaces of the two joining members where, by the increased interface resistance, the highest heat generation is obtained. After the welding temperature which is lower than the melting temperature of the weld metal – is reached, upset pressure is applied and the current circuit is opened. This produces a thick flash-free upset seam which is typical for this method. In order to guarantee the uniform heating of
the
abutting
faces, they must be conformable in their
before upset force has been applied
upset force
cross-sectional sizes and shapes
water-cooled clamping chucks (Cu electrodes)
with each other and they
must
have bulging at the end of the weld
parallel faces.
_ ~
As no molten metal br-er7-03e.cdr
develops
during
Process Principle of Pressure Butt Welding
pressure upset butt welding, the joining
Figure 7.3
surfaces must be free from contaminations
and
from
fixed clamping chuck
mobile clamping chuck
a+b b 2
oxides. Suitable for
clamping force
a
steel chuck
welding are unalloyed and low-alloy steels. The welding of aluminium and
copper shoe secondary side
copper material is, because of the tendency towards oxidation
and
primary side welding transformer br-er7-04e.cdr
Schematic Structure of a Flash Butt Welding Equipment
good
conductivity, possiFigure 7.4
a = flashing length b = upset loss
7. Pressure Welding
87
ble only up to a point. For the most part, smaller cross-sections with surfaces of up to 100 mm² are welded. Areas of applications are chain manufacturing and also extensions of wires in a wire drawing shop.
A flash butt welding equipment is, in its principal structure, similar to the pressure butt welding device, Figure 7.4.
While in pressure upset butt welding the
joining
members
are
always
strongly pressed together, in flash butt
br-er7-05e.cdr
© ISF 2002
welding only “fusing contact” is made during the heating phase. During the welding process, the workpiece ends
Figure 7.5
are progressively advanced towards each other until they make contact at several points and the current circuit is over these contact bridges closed. As the local current density at these points is high, the heating also develops very fast. The metal is liquified and, partly, evaporated. The metal vapour pressure causes the liquified metal to be thrown out of the gap. At the same time, the metal vapour is generating a shielding gas atmosphere; that is to say, with the exception of pipe welds, welding can be carried out without the use of shielding gas. The constant creation and destruction of the contact bridges causes the abutting faces to “burn”, starting from the contact points, with heavy spray-type ejection. Along with the occurrence of material loss, the parts are progressively advanced towards each other again. New contact bridges are created again and again. When the entire abutting face is uniformly fused, the two workpiece ends are, through a high axial force, abruptly pressed together and the welding current is switched off. This way, a narrow, sharp and, in contrast to friction welding, irregular weld edge is produced during the upsetting progress, which, if necessary, can be easy mechanically removed while the weld is still warm, Figure 7.5.
7. Pressure Welding
88
In flash butt welding, a fundamental distinction is made between two different working techniques. During hot flash butt welding a preheating operation precedes the actual flashing process, Figure 7.6. The preceding resistance heating is carried out by “reversing”, i.e., by the changing short-circuiting and pressing of the joining surfaces and by the mechanical separation in the reversed motion. When the joint ends are sufficiently heated, is the flashing process is initialised automatically and the following process is similar to cold flash butt welding. In contrast to cold flash butt welding, the advantage of hot flash butt welding is that, on one hand, sections of 20 times the size can be welded with the same machine efficiency and, on the other hand, a smaller temperature drop and with that a lower cooling rate in the workpiece can be obtained. This is of importance, especially with steels which because of their chemical composition have a tendency to harden. The cooling rate may also be reduced by conductive reheating inside the machine. A smooth and clean surface is not necessary with hot flash butt weld-
upset travel
ing. If the abutting
flashing travel
faces differ greatly from
the
desired
plane-parallelism,
upset force
an additional flashing process (a short
preheating
flashing
flashing
amperage
flashing period with time
low speed and high
hot flash welding
time
cold flash welding
br-er7-06e.cdr
energy) may be car-
Flashing Travel, Upset Travel, Upset Force and Welding Current in Timely Order
ried out first. Figure 7.6
The welding area of the structure of a flash butt weld shows a zone which is reduced in carbon and other alloying elements, Figure 7.7. Moreover, all flash butt welded joints have a pronounced coarse grain zone, whereby the toughness properties of the welded joint are lower than of the base metal. By the impact normalizing effect in the machine successive to the actual welding process, can the toughness properties be considerably increased. By one or several current impulses the weld
7. Pressure Welding
89
temperatures are increased by up to approximately 50° over the austeniting temperature of the metal. Steels, aluminium, nickel and copper alloys can be welded economically with the flash butt welding process. Supported heat affected zone
by the axial force,
10 mm
shrinkage in flash
material: C60 E
butt welding is so insignificant
that
0,1 mm
only very low residual stresses occur. It is, therefore, posweld
coarse grain zone
fine grain zone
soft-annealing zone
base metal
sible to weld also br-er7-07e.cdr
steels with a higher
Secondary Structure Along a Flash Butt Weld
carbon content. Figure 7.7
Profiles of all kind are butt welded with this method. The method is used n
for large-scale manufacture and with components of equal dimensions. The weldable cross-sections may reach dimensions of up to 120,000 mm². Areas of application are the welding of n
rails, the manufacture of car axles, F1 friction force
wheel rims and shafts, the welding of chain links and also the manufacture of tools and endless strips for pipe F2 upset force
production. Friction welding is a pressure welding method where the necessary heat
br-er7-08e.cdr
© ISF 2002
for joining is produced by mechanical friction. The friction is, as a rule, generated by a relative motion between a
Figure 7.8
7. Pressure Welding
90
rotating and a stationary workpiece while axial force is being applied at the same time, Figure 7.8.
After the joint surfaces are adequately heated, the relative motion is discontinued and the friction force is increased to upsetting force. An even, lip-shaped bead is produced which may be removed in the welding machine by an additional accessory unit. The bead is often considered as the first quality criterion.
Figure 7.9 shows all phases of the
br-er7-09e_sw.cdr
Phases of Friction Welding Process
friction welding process. In most cases this method is used for rota-
Figure 7.9 tion-symmetrical parts. It is, nowabrake
days, also possible to accurately join rectangular
and
polygonal
clamping tool
clamping tool
workpiece
clutch
pressure element for axial pressure
cross-
sections.
The most common variant of friction
conventional friction welding
welding is friction welding with a con-
driving motor
tinuous drive and friction welding with a flywheel drive, Figure 7.10. In fric-
flywheel
clamping tool
clamping tool workpiece
pressure element for axial pressure
tion welding with continuous drive, the clamped-on part to be joined is kept at a constant nominal speed by a drive, while the workpiece in the sliding chuck is pressed with a defined
flywheel friction welding
br-er7-10e.cdr
friction force. The nominal speed is maintained until the demanded temFigure 7.10
© ISF 2002
7. Pressure Welding
91
perature profile has been achieved. Then the motor is declutched and the relative motion is neutralised by external braking. In general, the friction force is raised to upsetting force after the rotation movement has been discontinued. During flywheel friction welding, the inertia mass is raised to nominal speed, the drive motor is declutched and the stationary workpiece is, with a defined axial force, pressed against the rotating workpiece. Welding is finished when the total kinetic energy - stored in the flywheel – has been consumed by the friction processes. This is the so-called self-breaking effect of the system. The method is used when, based on metallurgical processes, extremely short welding times may be realised. Further process variants are radial friction welding, orbital friction welding, oscillation friction welding and friction
surfacing.
However, process
friction welding time 1...100s
these variants
braking 0,1...0,5s
1800...
number of revolutions
5400 min
friction welding time 0,125...2s
900...
-1
-1
5400min
are until today still
time
in the experimental stage.
Recently,
axial pressure
40...280 20...100 Nmm
-2
40...280
-2
-2
Nmm
Nmm
new developments in the field of friction stud
welding
–
studs on plates –
torque
conventional friction welding
flywheel friction welding
br-er7-11e.cdr
have
been
intro-
Comparison of the Welding Processes for Conventional and Flywheel Friction Welding
duced. Figure 7.11
Figure 7.11 depicts the variation in time of the most important process parameters in friction welding with continuous drive and flywheel friction welding. The occuring moments’ maxima may be interpreted as follows: The first maximum, at the start of the frictional contact, is explained by the formation of local fusion zones and their shearing off in the lower temperature range. The torque decreases as a result of the risen temperature - which again is a consequence of the increased plasticity - and of the lowered deformation resistance. The second maximum is generated during the braking phase which precedes the spindle standstill. The second maximum is explained by the increased deformation resistance at dropping temperatures. The temperature drop in the joining zone is ex-
7. Pressure Welding
92
plained by the lowered energy input – due to the rotation-speed decrease – and also by the augmented radial displacement of the highly heated material into the weld upset.
In friction welding number of revolutions
with a continuous
upset force
drive, the process variation bined friction force
“comfriction
welding”
allows
the free and independent from each other selection of
reduction time
the
braking
and
© ISF 2002
br-er7-12e.cdr
upsetting
Combined Friction Welding
mo-
ments, Fig. 7.12. Figure 7.12 In this case, the rotation-energy which has been stored in the drive motor, the spindle and also in the clamping chuck, may be totally or partially converted by selfbreaking. Here, the breaking phase matches the breaking phase in flywheel welding. The use of this process variant allows the welding structures to influence each other in a positive way when many welding tasks are to be carried out. Moreover, the torque range may a)
P
b)
be accurately pre-
P
determined by the microcontroller
c)
the braking initiator
d)
P
P
of
P
P
which prevents the slip-through of the
e)
P
P
f)
workpieces in the
P
clamping chuck. © ISF 2002
br-er7-13e.cdr
Types of Friction Welding Processes
The universal friction welding ma-
Figure 7.13
7. Pressure Welding
93
chine is in its structure similar to a turning lathe, however, for the transmission of the high axial forces, the machine structure must be considerably more rigid. Basically, there are three types of friction welding: a) friction welding with a rotating workpiece and a translational motion of the other workpiece; b) friction welding with rotation and translational motion of one workpiece facing a stationary other workpiece, c) rotation and translation of two workpieces against a stationary intermediate piece. The remaining variations, shown in Figure 7.13, also find applications when both workpieces have to rotate in opposite direction to each other. For example, when a low diameter and, in connection with this, the low relative speeds demand the necessary heat quantity.
A survey of possible joint shapes achievable with friction welding is given in Figure 7.14. The specimen preparation of the joining members should, if possible, be carried out in a way that the heat input and the heat dissipation is equal for both members. Dependbefore welding
ing on the combina1. a)round stock with round stock
abutting
surfaces
b) round stock with round stock (different cross-sections, bevelled)
should be smooth, angular equal
and
of
3.
6. pipe with plate
0,75d
7. round material with plate, without preparation
d=0,75D
round stock with pipe
dimensions.
8. pipe with plate, without preparation (1/6)d © ISF 2002
br-er7-14e.cdr
A simple saw cut is,
Joint Types Obtained by Friction Welding
for many applications, sufficient.
d
d
The
5. round material with plate g/d » 0,25...0,3
d » 0,6D
2. a)round stock with round stock (different cross-sections, partially machined)
d
considerably.
after welding
g
tate the joining task
1..2°
D
b) round stock with round stock, chamfered
d
this provision facili-
before welding 4. pipe with pipe
D
tion of materials can
after welding
Figure 7.14
The method of heat generation causes a comparatively low joining temperature – lower than the melting temperature of the metals. This is the main reason why friction welding is the suitable method for metals and material combinations which are difficult to weld. It is also possible to weld material combinations (e.g. Cu/Al or Al/steel) which cannot be joined using other welding processes otherwise only with increased expenditure. Figure 7.15 shows a survey of possible material combinations. Many
7. Pressure Welding
94
combinations have, however, not yet been tested on their suitability to friction welding. Metallurgical reasons which may reduce the friction weldability are:
cirkon tungsten vanadium titanium tantalum stellite free cutting steel cast steel steel, austentic steel, high alloyed steel, alloyed steel, unalloyed silver niobium nickel alloys nickel molybdenum brass magnesium copper cobalt hard metal, sintered cast iron (GGG, GT) lead aluminium, sintered aluminium alloys aluminium
1. the quantity and distribution of
aluminium aluminium alloys aluminium, sintered lead cast iron (GGG, GT) hard metal, sintered cobalt copper magnesium brass molybdenum nickel nickel alloys niobium silver steel, unalloyed steel, alloyed steel, high alloyed steel, austentic cast steel free cutting steel stellite tantalum titanium vanadium tungsten cirkon
non-metal inclusions, 2. formation of low-melting or intermetallic phases, 3. embrittlement by gas absorption (as a rule, the costly, inert gas shielding can be dispensed with, even when welding titanium), 4. softening of hardened or pre-
friction weldable
cipitataly-hardened
restricted friction weldable not friction weldable
materials
and
not tested
5. hardening caused by too high br-er7-15e.cdr
a cooling rate.
© ISF 2002
By the adjustment of the welding paFigure 7.15
rameters in respect toweld joints, can
in most cases joints with good mechano-technological properties be obtained. The secondary structure along the friction-welded joint is depicted in Figure 7.16. An extremely grained (forge
finestructure structure) metal: S235JR
develops in the join10 mm
ing
zone
p = 30 N/mm2 t =6s 2 tSt = 250 N/mm n = 1500 U/min
region.
This structure which 1 mm
is typical of a fric-
structures on parallels with a 5 mm distance from the sample axis
tion-welded joint is characterised
by
high strength and
base metal
heat affected zone
transition heat affected zone - weld metal
br-er7-16e.cdr
toughness
proper-
Secondary Structure Along a Friction Weld
ties. Figure 7.16
weld metal
10 µm
7. Pressure Welding
95
Figure 7.17 shows a comparison between a flash butt-welded and a frictionwelded cardan shaft. The two welds are distinguished by the size of their heat affected zone and the development of the weld upset. While in friction welding a regular, lip-shaped upset is produced, the weld flash formation in flash butt welding is narrower and sharper and also considerably more irregular. Besides, the heat affected zone during friction welding is substantially smaller than during flash butt welding. Friction welding machines are fully mechanized and may well be integrated into production lines. Loading and unloading equipment, turning attachments for the preparation of the abutting surfaces and for upset removal and also storage units for complete welding programs make these machines well adaptable to automation. The machines
may
furthermore
be
equipped with parameter supervisory systems. During welding are parameters: welding path, pressure, rotational speed, and time are governed by the desired value/actual value comparison. This allows an indirect quality
flash butt welding
control. A further complement to the retension of parameters is the torque control, however this method is costly and it cannot be used for all applications because of its structural dimensions.
friction welding
br-er7-17e.cdr
© ISF 2002
Friction welding machines are mainly used in the series production and industrial mass production.
Figure 7.17 Nevertheless, these machines are also always applied when metals and material combinations which are difficult to weld have to be joined in a reliable and costeffective way. With the machines that are presently used in Germany, it is possible to weld massive workpieces in the diameter range of 0.6 up to 250 mm For steel pipes, the maximum weldable diameter is at present approximately 900 mm, the wall thicknesses are approx. 6 mm.
7. Pressure Welding
96
1
2
3
4
1 cardan shaft, AIZn 4,5 Mg 1 2 cardan shaft, retracted tube
1,2 joint ring
3 loading device
material combination: Cf53/ Ck45
4 unloading grippers
br-er7-18e_sw.cdr
3 cardan shaft, flattening test specimen 4 crown wheel, 16MnCr5/ 42Cr4 5 bimetal valve, X45CrSi9-3/ NiCr20 TiAl © ISF 2002
Figure 7.18
br-er7-19e_sw.cdr
© ISF 2002
Figure 7.19
Figures 7.18 to 7.20 show a selection of examples for the application of friction welding.
Figure 7.21 shows a comparison of the cost expenditure for the manufacture of a cardan shaft, carried out by forging and by friction welding, respectively. It shows that the application of the fric-
1 pump shaft 2 shaft C22E/ E295 3 press cylinder S185/9S 20K 4 hydraulic cylinder S235J3G2/ C60E or S235JR/ C15 5 cylinder case S235JR/ S355J2G3 6 piston rod 42Cr4 7 connecting rod 100Cr6/ S235JR 8 stud S235J2G3/ X5CrNi18-10 9 knotter hook 15CrNi6 br-er7-20e_sw.cdr
tion welding method may save approx. 20% of the production costs. This comparison is, however, not an universally © ISF 2002
valid statement as for each component a profitability evaluation must be carried out
Figure 7.20
7. Pressure Welding
97
separately. The comparison is just to show that, in many applications, considerable savings can be made if the matter of the joining technology by “friction welding” could be circulated to a wider audience of design and production engineers.
Figure 7.22 shows friction welds
in brief the impor-
160 mm
Ø40 mm
Ø30 mm
tant
advantages
and
disadvan-
tages of friction 940 mm
welding in comforged piece motor shaft
friction-welded piece € 20,-
flange,forged material costs shaft Ø30 und 40 mm 2x friction welds incl. upset removal
€ 20,-
€
7,50
€
4,25
€
3,-
€
14,75
parison
with
the
competitive method
of
flash
butt welding.
br-er7-21e.cdr
Cost Comparison of Forging/ Friction Welding in a Case of a Cardan Shaft
Figure 7.21 Pressure welding with magnetically impelled arc, “Magnetarc Welding”, is an arc pressure welding method for the joining of closed structural tubular shapes, Figure 7.23. The weldable wall thickness range is between 0.7 and 5 mm, the weldable diameter range between 5 and 300 mm. In “Magnetarc Welding” an arc burns between the joining surfaces and is rotated by external magnetic forces. This is achieved by a magnet coil system that produces a magnetic field.
The combined action of this magnetic
Advantages and disadvantages of friction welding in comparison with the competitive flash butt welding advantages: - clean and well controllable bulging - low heat influence on joining members - better control of heat input - low phase seperation phenomena in the bond zone - hot forming causes permanent recovery and recrystallisation processes in the welding area thus forming a very fine-grained structure with good toughness and strength properties (forged structure) - low susceptibility to defects, extremely good reproducibility within a wide parameter range - frequently shorter welding times - more choice in the selection of weldable materials and material combinations disadvantages: - torque-safe clamping necessary - machine-determined smaller maximum weldable cross-sections - susceptibility to non-metal inclusions - high expenditure requested because of high manufacturing tolerances - high capital investment for the machine br-er7-22e.cdr
field and the arc’s own magnetic field Figure 7.22
© ISF 2002
7. Pressure Welding
98
effects a tangential force to act upon the arc. The rotation of the arc heats and melts the joint surfaces. After an adequate heating operation, the two workpiece members are pressed and fused together. A regular weld upset develops which is normally not removed. The welding operation takes place under shielding gas (mainly CO2). 1. starting position
The shielding gas’
a) both workpieces are brought into contact b) welding current and magnetic field are switched on
function is not the
2. starting of welding
protection
a) both workpieces are seperated until a defined gap width is reached (retracting movement) - the arc ignites
weld from the sur-
of
rounding
3. heating
the
atmos-
phere but rather a
a) the arc rotates b) the joint surfaces are melting
contribution
to-
4. completion of welding
wards the stabilisa-
a) both workpieces are broght into contact again and upset b) welding current and magnetic field are switched off
tion of the arc. The
br-er7-23e.cdr
reproducibility Diagrammatic Representation of Magnetic Arc Welding
of
the arc ignition and
Figure 7.23
malleable
The prerequisite for the application of
cast steel
materials
free cutting steel
the weld bead are therefore improved.
steel, lowalloyed
steel, unalloyed
motion behaviour and the regularity of
a material in “Magnetarc Welding” is its steel, unalloyed
electrical conductivity and melting behaviour. Figure 7.24 gives a survey
steel, lowalloyed
of the material combinations which are
free cutting steel
nowadays already weldable under in-
cast steel
dustrial conditions. As reason is the symmetric heat input,
malleable
the subsequent upsetting of the liquid
suitable for magnetic arc welding
phase and the cooling off under pres-
not tested
sure. The cracking sensitivity of the br-er7-24e.cdr
welds is, in general, relatively low. This has a positive effect, particulary Figure 7.24
© ISF 2002
7. Pressure Welding
99
when steels with a high carbon content or machining steels are welded. The joining faces of the workpieces must be free from contamination, such as rust or scale. To obtain a defect-free weld, normally a simple saw cut is a sufficient preparation of the abutting surfaces.
If
special
demands are put on the dimensional accuracy
of
the
joints, the prefabrication have
tolerances to
be
ad-
justed accordingly. This applies also to © ISF 2002
br-er7-25e_sw.cdr
friction welding. Applications for Magnetic Arc Welding
Figure 7.25
Figures 7.25 and 7.26 show several application
examples
of
pressure
welding with magnetically impelled arc.
Figure 7.27 shows a summary of the most important advantages and disadvantages of this method in comparison with the competitive methods of friction welding and flash butt welding.
In friction-stir welding a cylindrical, mandrel-like tool carries out rotating self-movements between two plates which are knocked and clamped onto
br-er7-26e_sw.cdr
a fixed backing. The resulting friction heat softens the base metal, although Figure 7.26
© ISF 2002
7. Pressure Welding
100
the melting point is not reached. The plastified material is displaced by the Advantages and disadvantages of magnetic arc welding in comparison with flash butt welding and/ or friction welding
mandrel and transported behind the tool where a longitudinal seam devel-
advantages:
ops.
- lower energy demands - material savings through lower loss of length - better dimensional accuracy in joining especially for small
The advantages of this method which
wall thicknesses - in comparison with friction welding less moving parts (only axial movement of one joining member during upsetting)
is mainly used for welding of aluminium alloys is the low thermal stress of
- no restrictions to the free clamping length - smaller and more regular welding edge
the component which allows joining
- no spatter formation
with a minimum of distortion and
disadvantages: - suitable for small wall thicknesses only
shrinkage. Welding fumes do not de-
(maximum wall thickness: 4 - 5 mm)
velop and the addition of filler metal or
- welding parameters must be kept within narrow limits - only magnetizing steels are weldable without any difficulties
shielding gases is not required. br-er7-27e.cdr
© ISF 2002
Figure 7.27
workpiece tool collar
fixed backing
contoured pin
br-er7-28e.cdr
Friction-Stir Welding
Figure 7.28
8. Resistance Spot Welding, Resistance Projection Welding and Resistance Seam Welding
2003
8. Resistance Spot-, Resistance Projection- and Resistance Seam Welding
101
Figure 8.1 shows an extract from the classification of the welding methods according to DIN 1910 with a detailed account of the conductive resistance pressure welding.
In the case of resistance pressure welding, the heating occurs at the welding point as a consequence of Joule resistance heating caused by current flow through an electrical conductor, Figure 8.2. In spot and projection welding, the plates to be welded in overlap. Current supply is carried out through spherical or flat electrodes, respectively. In roller seam welding, two driven roller electrodes are applied. The plates to be
welded
are
mainly overlapped.
welding
The heat input rate pressure welding
fusion welding
Qinput is generated by resistance heat-
cold pressure welding
resistance pressure welding
induction pressure welding
Conduction pressure welding
resistance spot welding
projection welding
roller seam welding
friction welding
ing in a currentcarrying conductor, Figure 8.3. How-
resistance butt welding
flash butt welding
ever, only the effective heat quan-
© ISF 2002
br-er8-01e.cdr
tity Qeff is instru-
Classification of Welding According to DIN 1910
mental in the formation of the weld
Figure 8.1
nugget. Qeff is comspot welding
roller seam welding
projection welding
l workpieces overlap l electrode l weld nugget
l workpiece usually in general overlap l driven roller electrode l spot rows (stitch weld, roller spots)
l workpieces with elevations (concentration of electicity) l workpieces overlap l pad electrode l several joints in a single weld l weld nugget joint
posed of the input
1
2
1
3
2
2
1
3
heat minus the dissipation heat. The heat
loss
arises
from
the
heat
3
dissipation into the 4 5
1
1
plates
1
1 electrode force 2 elektrodes 3 production part br-er8-02e.cdr
Figure 8.2
4 loaded area
electrodes and the
5 projection
and
also
from thermal radiation.
8. Resistance Spot-, Resistance Projection- and Resistance Seam Welding
102
The resistance during resistance heating is composed of the contact resistances on the two plates and of their material resistance. The reduction of the electrode force down to 90% increases the heat
electrode force effective heat total heat input current (time dependence) heat losses losses into the electrodes losses into the sheet metal losses by heat radiation total resistance material resistance contact resistance
Q4
Q2
rate
by
105%, the reduc-
Q4
tion of the welding Q3
Qeff = Qinput - Q1l
Qeff
Q3
Q4
t=tS
Qinput = C
input
Fel
Q4
down
to
90%
decreases
Q2
the heat rate to
Fel
80% and a welding
2
I (t) R(t) dt
t=0
current
Q1 = Q2 + Q3 + Q4
time reduction to
R(t) = Rmaterial(t) + Rc (t) br-er8-03e.cdr
90%
decreases
the heat rate to 92%.
Figure 8.3
The time progression of the resistance is shown in Figure 8.4. The contact resistance is composed of the interface resistances between the electrode and the plate (electrode/plate) and between the plates (plate/plate). The resistance height is greatly dependent on the applied electrode force. The higher this force is set, the larger are the conductive
cross-
theoretical contact area 100% metallic conduction contact
sections
proportion at room temperature
contact points and
at
the
mOhm
resistance
total resistance
low electrode force high resistance
tances. The con-
sum of material resistance
high electrode force low resistance proportion after first milliseconds welding time
sum of contact resistances
5
10
welding time
smaller the resis-
periods
surface resistance is collapsed, a3 is highly extended A1: area out-of-contact A2: contact area with high resistance A3: contact area completely conductive
tact
surfaces,
which are rapidly increasing at the start
of
welding,
effect a rapid re-
br-er8-04e.cdr
duction of interface resistances. Figure 8.4
8. Resistance Spot-, Resistance Projection- and Resistance Seam Welding
103
With the formation of the weld nugget the interface resistances between the plates disappear. During the progress of the weld the material resistance increases from a low value (surrounding temperatures) to a maximum value above the melting temperature.
Figure 8.5 shows diagrammatically the different resistances during the spot welding process with acting electrode force, but without welding current. Weld nugget formation must therefore start in the joining zone because of the existing high contact resistance there. electrode force
resistance rate
Figure 8.6 shows R1
directly cooled elec-
R3
R3
trodes
for
resis-
R6 R6
_ ~
tance welding. The
R7 R4
coolant is normally
R5
the
R7
R2
water. In the cooling tube,
R5
R4 0
cooling
100
200
R [µOhm]
water is transported to
the
electrode
br-er8-05e.cdr
base. The diagram shows the temperature distribution in
Figure 8.5
the electrodes and cooling tube
in the plates. The 6-8
maximum tempera-
cooling drill-hole
2-5
10 - 20
slope
ture is reached in the centre of the weld
nugget
and
decreases strongly in the electrode di-
°C
rection. © ISF 2002
br-er8-06e.cdr
Electrode Cooling
Figure 8.6
8. Resistance Spot-, Resistance Projection- and Resistance Seam Welding
104
Sequence of a resistance spot welding process, Figure 8.7: 1 ->2 Lowering of the top electrode 2->3 Application of the adjusted electrode force Set-up time tpre, sequence 3->4 Switching-on of the adjusted welding current for the period of the welding time tw. Formation of the weld nugget in the joining zone of both workpieces. An example shows the macrosection of a weld nugget after the welding time has ended. 4->5 Maintaining the electrode force for the period of the set post-weld holding time th. 5->6 Switching-off the force generating system and lifting the electrodes off the workpiece.
The functions of the set-up time and the post-weld holding time are listed in Figure 8.8. Dependent on the welding task different force and current programs can be set in the welding machines, Figure 8.9. In the top the simplest possible welding program sequence is shown: The application of the electrode force, the set-up time sequence tpre, the switching-on of the welding current and the sequence of the weldFel
Iw
set-up time
electrode force Fel
- compressing the workpiece - build-up of electrode force to preset value - setting-up of reproducible resistance before welding - electrode resting after bounce - preventing resting of electrode on workpiece under electricle voltage
welding current Iw
time t
tpre
tw
th
top electrode
postweld-holding time - holding time of workpiece during cooling of molten metal - prevention of pore formation in the welding nugget - prevention of lifting the electrode under voltage
workpiece lower electrode
insufficiently melted weld nugget
weld nugget
The postweld-holding time has influence on the weld point hardening within certain limits.
totally melted weld nugget
br-er8-07e.cdr
© ISF 2002
br-er8-08e.cdr
Time Sequence of Resistance Spot Welding
Figure 8.7
© ISF 2002
Functions of Pre- and Postwelding
Figure 8.8
8. Resistance Spot-, Resistance Projection- and Resistance Seam Welding
105
welding current
electrode force
ing time tw, the sequence of the postFel
tw tpres
th
time
Fel
welding current
electrode force electrode force
tpre = pre-weld time tw = welding time th = holding time tpres = pressure time
weld holding time th and the switchingoff of the force generating system. The diagram in the centre is almost identi-
tpre
welding current
Iw
5
cal to the one just described.
Merely in the welding current range,
Iw
5 7
welding is carried out using an adjust-
8
able current rise (7) and current decay
1 2 5 7
Fel
3 Iw 6
4
8
time
(8). The diagram below depicts a more
1 - initial force 2 - welding pressure force 3 - post pressure force 4 - preheating current 5 - welding current 6 - postheating current 7 - ascending current 8 - descending current
sophisticated current program. In addi-
time br-er8-09e.cdr
tion, welding is carried out with a variable electrode force (2) and with preheating (4) and post-heating current (6). Dependent on the control system,
Course of Force and Current
the process can be influenced by adjustment.
Figure 8.9 A controlled variable may be, for in1
stance, the electrode path, the resistance progress, the welding current or
9
the welding voltage.
10
2 3
6
Figure 8.10 shows the principle struc-
11
12 4
ture of a resistance spot welding
7
machine. The main components are:
5 8
the machine frame, the welding transformer with secondary lines, the elec-
1 electrode force cylinder 2 pneumatic equipment 3 machine tool frame 4 welding transformer 5 power control unit 6 current conductor 7 lower arm 8 foot switch 9 top arm 10 electrical power supply cable 11 water cooled electrode holder 12 electrode
trode pressure system and the control system. This principle design applies to spot, projection and roller seam welding machines. Differences are to be
br-er8-10e.cdr
found merely in the type of electrode
© ISF 2002
Schematic Assembly of Spot Welding Machine
fittings and in the electrode shapes. Figure 8.10
8. Resistance Spot-, Resistance Projection- and Resistance Seam Welding
106
Figure 8.11 depicts the possible process variations of resistance spot welding. These are distinguished by the number of plates to be welded and by the arrangement of the electrodes or, respectively, of the transformers. It has to be noted that with a corresponding arrangement also plates can be welded where one of the two plates has a non-conductive surface (as, for example, plastics). Figure 8.12 shows the current types which are normally used for resistance welding. Alternating current has the simplest structure (Figure 8.13) and is most price effective, unavoidable are, however, the disadvantages of current zeros and weld nugget cooling. In relation to the average cur-
~
rent values, peak ~
loads
occur
~
and,
with that, increased electrode These
wear.
two-sided single-shear single-spot welding
two-sided two-shear spot welding (stack welding)
one-sided single-spot welding with contact electrode
~
~
~
extreme +
peak loads do not with
direct
+
+
+
occur
+
current.
~
two-sided duplex spot welding
one-sided duplex spot welding with conductive base
one-sided multi-spot welding with conductive base © ISF 2002
br-er8-11e.cdr
The structural de-
Variants of Spot Welding
sign of a d.c. supply unit
is,
more
however,
Figur 8.11
complicated
alternating current
medium frequency direct current
expensive
than an a.c. supply
12
[kA]
15 10 5 0 0.00 0.02 0.04 0.07 0.09 0.11 0.13 0.16 -5
current
more
therefore,
current
and,
[kA]
20
-10 -15 -20
6 4 2 0.00 0.02 0.04 0.06 0.08 0.10 0.12 0.14 0.16
welding time [s]
supply, the welding
[kA]
operate with a 50
18 16 14 12 10 8 6 4 2 0 0.00
0.02 0.04
0.06 0.08 0.10 0.12
welding time
0.14 0.16
45 40 35 30 25 20 15 10 5 0 0.00
Figur 8.12
0.06 0.08 0.10
0.12 0.14 0.16
[s] © ISF 2002
Current Types
trolled only in 20 ms
0.02 0.04
welding time
[s]
br-er8-12e.cdr
current can be con-
[s]
impulse capacitor current
current
machines
[kA]
conventional
Hz primary current
welding time
"conventional" direct current
current
welding
8
0
unit. As
10
8. Resistance Spot-, Resistance Projection- and Resistance Seam Welding
units single-phase alternating current
static-inverter direct current
107
(1
period).
When the inverterdirect current technique or, respectively, the medium-
3-phase direct current
frequency
capacitor impulse discharge
tech-
nique is used, a finer setting of the current-on
period
and a more precise br-er8-13e.cdr
control of the welding current is posFigure 8.13
sible.
In order to realise higher currents and shorter welding times, the impulse capacitor resistance welding technique is applied. The rectified primary current is stored in capacitors and, through a high-voltage transformer, converted to high welding currents. The advantages of this technique are low heat input and high reproducibility. Because of the high energy density, materials with good conductivity can be welded and also multiple-projection welds can be carried out. A disadvan-
electrodes form A
form B
form C
form E
form F
form G
form D
tage of this method is, apart from the high equipment costs, the difficult regulation of the welding current. electrode caps
Electrodes for spot resistance welding have the property of transferring the electrode force and the welding current. They are wearing parts and, therefore, easily replaceable. Depend-
electrode holders br-er8-14e.cdr
ing on the shape and type of elec-
© ISF 2002
Electrodes, Electrode Caps and Holders
trode, solid electrodes or electrode Figure 8.14
8. Resistance Spot-, Resistance Projection- and Resistance Seam Welding
108
caps, must be either remachined or recycled. Figure 8.14 depicts various
requirements - good electrical conductivity - good thermal conductivity - high high-temperature strength - high temperature stability - high softening temperature - little tendency to alloying with workpiece material - easy options in machining
ISO 5182 Group
Type 1 2
A
3
4
Group Type
No. 1
Cu - ETP
2
Cu Cdl
1
Cu Crl
2
Cu Crl Zr
1
Cu CO2 Be
2
Cu Ni2 Si
1
Cu Ni1 P
2
Cu Be2 Co Ni
3
Cu Ag6
4
CuAl10NiFe5Ni5
and holders.
Dependent upon the electrode application, different alloyed electrode ma-
ISO 5182
Key
types of electrodes, electrode caps
Key
terials are used, Figure 8.15. The
No.
10
W75 Cu
11
W78 Cu
12
WC70 Cu
red hardness, the tempering resis-
13
Mo
tance, the conductivity, the fusion
14
W
15
W65 Ag
added alloying elements influence the
B
temperature, the electrode alloying tendency, and, finally, the machinability of the electrode material. When
br-er8-15e.cdr
beryllium is used as an alloying eleElectrode Materials
ment, the admissible MAC values
Figure 8.15 poor
good
must be strictly adhered to during remachining or dressing of the electrodes.
Already during the design phase of the components to be welded, importance must be attached to a good accessibility of the welding point. Moreover, the electrode force which is imperative to the process must be applied in a way that no damage is done to the workpiece. In the ideal case, the welding point is accessible from the top and from below, Figure 8.16.
br-er8-16e.cdr
© ISF 2002
Accessibility for Spot Welding Electrode
Figure 8.16
8. Resistance Spot-, Resistance Projection- and Resistance Seam Welding
poor
109
In order to avoid the displacement of
good
the electrodes, the electrode working surface must be flat. Also during the design phase space must be provided for an adequately large clearing zone around the working point, in order to guarantee the unimpeded electrode approach to the working point, Figure 8.17.
Dependent on the joining job, the process variation, or the resistance welding method, a so-called “shunt current/effect” may be noticed. This br-er8-17e.cdr
© ISF 2002
Contact Area for Spot Welding Electrodes
current component, as a rule, does not contribute to the formation of the
Figure 8.17 weld nugget; under certain circumstances it might even prevent a reliable welding process. In the example, shown in Figure 8.18, the shunt cur-
spot welding
rent leads to undesired fusing contacts
A
and, because of the lacking electrode
shunt connection current
force at this point, also to damages to the plate surface. copper
current path
indirect welding one side
If unsuitable welding parameters have been set, weld spatter formation may occur, Figure 8.19. Liquid molten metal forms on the plate surface or in
roller seam welding
the joining zone. The reasons for this
br-er8-18e.cdr
kind of process disturbance are, for
Shunting
Figure 8.18
8. Resistance Spot-, Resistance Projection- and Resistance Seam Welding
110
example, too low an electrode force Welding spatter: Discharge of molten material between two steel sheets or from the surface of steel sheets.
with regard to the set welding current or welding time, too high an energy input with regard to the plate thickness or too small an edge distance of the welding point.
fig. 1
Figure 8.20 shows a list of a large
fig. 2
number of possible disturbances in
Reason here is high welding current, (fig. 1) or too-small edge distance (fig. 2)
resistance spot welding. Welding current changes are caused by: shunt, electrode wear, cable wear, mains voltage variations, secondary
porosity in the joint caused by welding spatter
discharge of molten material at the joint plane
impedance.
br-er8-19e.cdr
Welding Spatter
Figure 8.19
Different welding conditions are caused by welding machine wear, different heat dissipation. Non-uniform conditions by alterations to components are: different plate plate
quality, number of plate
sur-
faces,
edge
dis-
tances.
Electrode
force changes are caused sure
by:
pres-
shunt connection
alteration to force
plates,
welding current changes
alteration of pressure
wear of electrodes
wear of cable
mains voltage fluctuation
secondary electrical impedance
Qeff = Qinput - Qlosses
wear Qeff diversion heat
plate
fluctuations
and -changes, plate
plate thickness
bouncing.
quality of plates
number of plates
plate surface
modification of the unit br-er8-20e.cdr
Figure 8.20
edge distance
welding equipment
thicknesses,
8. Resistance Spot-, Resistance Projection- and Resistance Seam Welding
111
The resistance spot welding method allows welding of a large number of weldability
materials
aluminium
alloying elements
good weldability
sufficient weldability
satisfactory maximum content [%]
iron
very good
gold
satisfactory
C
0,25
0,40
C + Cr
0,35
1,60
C + Mo
0,50
0,70
C+V
0,40
0,60
C + Mn
1,40
1,60
molybdenum satisfactory
C + Ni
3,00
4,00
nickel
very good
Si
0,40
1,00
platinum
very good
Cu
0,60
0,60
P+S
0,10
0,10
C+Cr+Mo+V
0,60
1,60
cobalt
very good
copper
poor
magnesium
good
silver
very good
tantalum
very good
titanium
very good
tungsten
satisfactory
materials. A list of the various materials is shown in Figure 8.21. The alloying elements which are used for steels have a varying influence on the suitability for resistance spot welding. The values which are indicated in the table are valid only when the stated element is the sole alloying constituent of the steel material.
influence of alloying elements (steel materials)
weldable materials br-er8-21e.cdr
Figure 8.22 shows a comparison between resistance spot and resistance projection welding. The fun-
Weldable Materials
damental difference between the two methods lies in the definition of the
Figure 8.21
current transition point.
The differences between both methods are illustrated in Figure 8.23. The short life of the electrodes used for resistance spot welding is explained by the higher thermal load and the larger pressing area caused by the smaller electrode contact areas. The term “electrode life” stands for the num-
after welding
before welding
ber of welds that can be carried out with
one
pair
electrodes
of
without follow-up distance
further rework and without
exceeding elektrode
the tolerances for quality criteria of the
projection br-er8-22e.cdr
weld.
Figure 8.22
8. Resistance Spot-, Resistance Projection- and Resistance Seam Welding
spot welding
projection welding
up to 20 mm
> 20 mm
embossed projection shape
elektrodes: diameter tip face
pressed mould pressed
convex
flat
electrode life
less
longer
place where the nugget originates
elektrodes
projections
one
several
small
big
current distribution
no
yes
force distribution
no
yes
number of welding nuggets
circular longitudinal annular
solid projection shape
112
natural projection shape
struck machined cut pushed
circular longitudinal annular interrupted annular
spot contact line contact
Circular
follow-up distance
weld nut
problems:
br-er8-23e.cdr
© ISF 2002
Longitudinal
cut
Annular
pushed
crossed wires
wire-plate
bolt-pipe
br-er8-24e.cdr
Differences Between Resistance Spot and Projection Welding
Customary Projection Shapes
Figure 8.23
Figure 8.24
die
die plate
Depending on the demands on the joint strength or on the projection rigidity, dif-
plate
ferent projection shapes are applied. These are annular, circular or longitudid1
mould plate
mould plate
counter-die
nal projections. The welding projections
d1
are, according to their size, adapted to ring projection
embossed projection
the used plate thickness and may, therefore, appear as different types in the
die
workpiece: embossed projections, solid
b
l mould plate
plate
projections and natural projections. The survey is shown in Figure 8.24.
longitudinal projection br-er8-25e.cdr
© ISF 2002
Production of Embossed Projection Shapes
Figure 8.25
In Figure 8.25 the production of embossed projections in different shapes is shown. The shape is embossed onto the
8. Resistance Spot-, Resistance Projection- and Resistance Seam Welding
113
plate surface by appropriate die plates, dies and, if necessary, counter dies.
alternating current distribution intensity of current increases from the center to the outer area caused by current displacement
force distribution of a C-frame projection press welder during bending of machine tool frame
direct current distribution intensity of current decreases from the center to the outer area caused by the longer current path
force distribution of a C-frame projection press welder with non-parallel positioning tables
br-er8-26e.cdr
© ISF 2002
br-er8-27e.cdr
© ISF 2002
Problem of Current Distribution During Projection Welding
Figure 8.26
Problems of Force Distribution During Projection Welding
Figure 8.27
Various problems occur in projection welding caused by the welding of several joints in a single working cycle. Due to different current paths - when using direct current - and a current displacement - when lap joint
using alternating cur-
lap joint with wire electrode
lap joint with foil
squash seam weld
butt weld with foil
rent -, welding nuggets
with
qualities
differing are
pro-
before welding
duced when no preventive remedies are taken, Figure 8.26.
after welding
A varying force distri© ISF 2002
br-er8-28e.cdr
bution, as shown by Roller Seam Welding
the example in Figure Figure 8.28
8. Resistance Spot-, Resistance Projection- and Resistance Seam Welding
114
8.27, also leads to differing qualities of the produced weld nuggets.
In Figure 8.28 several examples of application using projection welding
interrupted-current roller seam weld
are depicted. In this example, the shapes are of the embossed type.
Figures 8.29 and 8.30 show several overlap seal weld
process variations of roller seam welding. Seam welding is actually a continuous spot welding process, but with the application of roller electrodes. In contrast to resistance spot
continuous D.C. seal weld br-er8-29e.cdr
© ISF 2002
Weld Types for Roller Seam Welding
welding the electrodes remain in contact and turn continuously after the first weld spot has been produced. At
Figure 8.29
the points where a welding spot is to
be produced again current flow is initiated. Dependent on the electrode feed rate
and
on
the
welding current frequency, spot welds or seal welds with overlapping
weld
nuggets
pro-
are
duced. The applica-
br-er8-30e_sw.cdr
tion of d.c. current also produces seal welds.
© ISF 2002
Application Examples of Projection Welding
Figure 8.30
9. Electron Beam Welding
2003
9. Electron Beam Welding
115
The application of highly accelerated electrons as a tool for material processing in the fusion, drilling and welding process and also for surface treatment has
high voltage supply
been known since the Fifties. Ever control elektrode
since, the electron beam welding
anode
process has been developed from the
adjustment coil
laboratory stage for particular applica-
to vacuum pump
valve
viewing optics
working chamber
beam forming and guidance
beam generation
cathode
stigmator
tions. In this cases, this materials could not have been joined by any
focussing coil defelction coil
industrially
applied
high-production
joining method. The electron beam welding machine
workpiece workpiece handling
to vacuum pump
is made up of three main components: beam generation, beam manipulation
chamber door br-er9-01e.cdr
© ISF 2002
Schematic Representation of an Electron Beam Welding Machine
and forming and working chamber. These components may also have separate vacuum systems, Figure 9.1.
Figure 9.1
power supply
A tungsten cathode which has been heated under vacuum emits electrons by thermal emission. The heating of
chamber evacuation system valve
evacuation system for gun control cabinet EB-gun
the tungsten cathode may be carried out directly - by filament current - or indirectly - as, for example, by coiled filaments. The electrons are accelerated by high voltage between the cathode and the pierced anode. A modulating electrode, the so-called “Wehnelt cylinder”, which is positioned
working chamber workpiece receiving platform workpiece handling
between anode and cathode, regulates the electron flow. Dependent on
control panel control desk
br-er9-02e_f.cdr
© ISF 2002
All-Purpose EBW Machine and Equipment
the height of the cut-off voltage beFigure 9.2
9. Electron Beam Welding
116
tween the cathode and the modulating electrode, is a barrier field which may pass only a certain quantity of electrons. This happens during an electron excess in front of the cathode where it culminates in form of an electron cloud. Due to its particular shape which can be compared to a concave mirror as used in light optic, the Wehnelt cylinder also effects, besides the beam current adjustment, the electrostatic focussing of the electron beam. The electron beam which diverges after having passed the pierced anode, however, obtains the power density which is necessary for welding only after having passed the adjacent alignment and focussing system. One or several electromagnetic focussing lenses bundle the beam onto the workpiece inside the vacuum chamber. A deflection coil assists in maintaining the electron beam oscillating motion. An additional stigmator coil may help to correct aberrations of the lenses. A viewing optic or a video system allows the exact positioning of the electron beam onto the weld groove.
The core piece of the electron beam welding machine is the electron beam gun where the electron beam is generated under high vacuum. The tightly focussed electron beam diverges rapidly under atmospheric pressure caused by scattering and ionisation development with air. As it would, here, loose power density and efficiency, the welding process is, as a rule, carried out under medium or high vacuum. The necessary vacuum is generated in separate vacuum pumps for working chamber and beam gun. A shut-off valve which is positioned between electron gun and working chamber serves to maintain the gun vacuum while the working chamber is flooded. In universal machines, Figure
9.2,
the
back-scattered electrons
x-ray
workpiece maniputhermal radiation
lator assembly inside
the
secondary electrons
vacuum
chamber is a slide x
convection
with working table positioned over NCcontrolled
stepper
y
heat conduction
motors. For work-
z © ISF 2002
br-er9-03e.cdr
piece removal, the Energy Transformation Inside Workpiece
slide is moved from Figure 9.3
9. Electron Beam Welding
117
the vacuum chamber onto the workpiece platform. A distinction is made between electron beam machines with vertical and horizontal beam manipulation systems.
The energy conversion in the workpiece, which is schematically shown in Figure 9.3, indicates that the kinetic energy of the highly accelerated electrons is, at the operational point, not only converted into the heat necessary for welding, but is also released by heat radiation and heat dissipation. Furthermore, a part of the incident electrons (primary electrons) is subject to backscatter and by secondary processes the secondary electrons are emitted from the workpiece thus generating X-rays.
The impact of the electrons, which are tightly focussed into a corpuscular beam, onto the workpiece surface stops the electrons; their penetration depth into the workpiece is very low, just a few µm. Most of the kinetic energy is released in the form of heat. The high energy density at the impact point causes the metal to evaporate thus allowing the following electrons a deeper penetration. This finally leads to a metal vapour pour cavity cavity which which is is surrounded by a shell of fluid metal, covering the entire weld surrounded
by
a
shell of fluid metal, covering the entire weld depth, Figure 9.4. This deep-weld effect allows nowadays
penetration
depths
into
steel
a)
b)
c)
d)
materials of up to 300
mm,
when
Principle of Deep Penetration Welding
modern high vacuum-high
voltage
© ISF 2002
br-er9-04e.cdr
Figure 9.4
machines are used.
The diameter of the cavity corresponds approximately with the beam diameter. By a relative motion in the direction of the weld groove between workpiece and electron beam the cavity penetrates through the material, Figure 9.5. At the front side of the cavity new material is molten which, to some extent, evaporates, but for the most part
9. Electron Beam Welding
118
flows around the cavity and rapidly solidifies at the backside. In order to maintain the welding cavity open, the vapour pressure must press the molten metal round the vapour column against the cavity walls, by counteracting its hydrostatic pressure and the surface tension.
However, this equilibrium of forces is unstable. The transient pressure and temperature conditions inside the cavity as well as their respective, momentary diameters are subject to dynamic changes. Under the influence of the resulting, dynamically changing geometry of the vapour cavity and electron beam
motion of the molten metal groove
groove front side
keyhole
melting pool
molten zone
welding direction
vapour capillary
with an unfavourable
selection
the
welding
rameters,
F1
of pa-
metal
fume bubbles may F2
solidified zone
be included which
F3
on cooling turn into
F1
F1 : force resulting from vapour pressure F2 : force resulting from surface tension F3 : force resulting from hydrostatic pressure
shrinkholes, Figure © ISF 2002
br-er9-05e.cdr
9.6. The unstable pres-
Condition in Capillary
sure exposes the Figure 9.5
molten backside of the vapour cavity to strong and irregular changes in shape (case II). Pressure variations interfere with
β
I
II
III
the
regular
flow at the cavity backside, act upon the molten metal
workpiece movement © ISF 2002
br-er9-06e.cdr
Model of Shrinkage Cavity Formation
and, in the most unfavourable case, press the unevenly
Figure 9.6
9. Electron Beam Welding
119 distributed groove
metal into different
g len
t
fs ho
zones of the mol-
m ea
len
o gth
ea fs
m
ten
blind bead
back-
the so-called vapour pockets. The
molten area
cavities
unapproachable gap lower bead
cavity
side, thus forming
weld thickness
weld penetration depth
width of seam Nahtdicke
weld reinforcement
upper bead
root reinforcement
molten
end crater
are
not
always filling with
root weld
molten metal, they © ISF 2002
br-er9-07e.cdr
collapse
Basic Definitions
sporadi-
cally and remain as Figure 9.7
hollow spaces after
solidification (case Ill). The angle ß (case I) increases with the rising weld speed and this is defined as a turbulent process. Flaws such as a constantly open vapour cavity and subsequent continuous weld solidification could be avoided by selection of jobsuitable welding parameter combination and in particular of beam oscillation characteristics, it has to be seen to a constantly of the molten metal, in
by accelerating voltage: l high voltage machine (UB=150 kV) l low voltage machine (UB=60 kV)
order to avoid the above-mentioned defects. Customary beam oscillation types are: circular, sine, double parab-
by pressure: l high vacuum machine l fine vacuum machine l atmospheric machine (NV-EB welding)
ola or triangular functions.
Thick plate welding accentuates the process-specific
advantage
of
the
by machine concept: l conveyor machine l clock system l all-purpose EBW machine l local vacuum machine l mobile vacuum machine l micro and fine welding machine
deep-weld effect and, with that, the possibility to join in a single working cycle with high weld speed and low heat input quantity. A comparison with br-er9-20e.cdr
the submerged-arc and the gas metalClassification of EBW Machines
arc welding processes illustrates the depth-to-width ratio which is obtainFigure 9.8
9. Electron Beam Welding
120
able with the electron beam technology, Figure 9.7. Electron beam welding of thick plates offers thereby decisive advantages. With modern equipment, wall thicknesses of up to 300 mm with length-to-width ratios of up to 50 : 1 and consisting of low and high-alloy materials can be welded fast and precisely in one pass and without adding any filler metal. A corresponding quantification shows the advantage in regard of the applied filler metal and of the primary energy demand. Compared with the gas-shielded narrow gap welding process, the production time can be reduced by the factor of approx. 20 to 50.
Numerous specific advantages speak in favour of the increased application of this high productivity process in the manufacturing practice, Figure 9.8. Pointing to series production, the high profitability of this process is dominant. This process depends on highly energetic efficiency -6
< 1 x 10 mbar
together with a sparing use of resources during fabrication.
< 5 x 10-4 mbar
Considering the above-mentioned advantages, there are also disadvantages which emerge from the process. These are, in particular, the high cooling rate, the high equipment costs and the size of the chamber, Figure 9.9.
In accordance with DIN 32511 (terms for methods and equipment applied in electron and laser beam welding), the
br-er9-09e_f.cdr
specific designations, shown in Figure
EB-Welding in High Vacuum
9.10, have been standardised for electron beam welding.
Figure 9.9
Electron beam units are not only distinguished by their working vacuum quality or the unit concept but also by the acceleration voltage level, Figure 9.11. The latter exerts a considerable influence onto the obtainable welding results. With the increasing acceleration voltage, the achievable weld depth and the depth-to-width ratio of the weld
9. Electron Beam Welding
121 geometry are also increasing. A disadvantage of the increasing accelerating voltage is, however, the exponential increase of X-rays and, also, the
< 1 x 10-6 mbar
likewise increased sensitivity to flashover voltages. In correspondence with
-2
< 5 x 10 mbar
the size of the workpiece to be welded and the size of the chamber volume, high-voltage beam generators (150 200 kV) with powers of up to 200 kW are applied in industrial production, while
the
low-voltage
technology
(max. 60 kV) is a good alternative for smaller units and weld thicknesses. br-er9-10e_f.cdr
The design of the unit for the lowEB-Welding in Fine Vacuum
voltage technique is simpler as, due
Figure 9.10 to the lower acceleration voltage, a separate complete lead covering of the unit is not necessary. While during the beam generation, the vacuum (p = 10-5 mbar) for the insula-
-4
< 1x 10 mbar
tion of the beam generation compart-1
~ 10 mbar
ment and the prevention of cathode
~ 1 mbar
oxidation is imperative, the possible working pressures inside the vacuum chamber vary between a high vacuum (p = 10-4 mbar) and atmospheric pressure. A collision of the electrodes with the residual gas molecules and the scattering of the electron beam which is connected to this is, naturally, lowest
br-er9-11e_f.cdr
Atmospheric Welding (NV-EBW)
in high vacuum. Figure 9.11
9. Electron Beam Welding
122 The beam diameter is minimal in high vacuum and the beam power density
in vacuum
is maximum in high vacuum, Figure
r
thin and thick plate welding (0,1 mm bis 300 mm)
r
extremely narrow seams (t:b = 50:1)
r
low overall heat input => low distortion => welding of completely processed components
r
high welding speed possible
r
no shielding gas required
(narrow, deep welds with a minimum
r
high process and plant efficiency
energy input) or the choice of the ma-
r
material dependence, often the only welding method
terials to be welded (materials with a
9.12. The reasons for the application of a high vacuum unit are, among others, special demands on the weld
high oxygen affinity). The application
at atmosphere r
very high welding velocity
of the electron beam welding process
r
good gap bridging
also entails advantages as far as the
r
no problems with reflection during energy entry into workpiece
structural design of the components is concerned.
br-er9-12e.cdr
© ISF 2002
Advantages of EBW
With a low risk of oxidation and reduced demands on the welds, the so-
Figure 9.12 called “medium-vacuum units” (p = 102
mbar) are applied. This is mainly be-
in vacuum r
electrical conductivity of materials is required
r
high cooling rates => hardening => cracks
for instance, the reduction of cycle
r
high precision of seam preparation
times, Figure 9.13. Areas of applica-
r
beam may be deflected by magnetism
tion are in the automotive industry (pis-
r
X-ray formation
tons, valves, torque converters, gear
r
size of workpiece limited by chamber size
r
high investment
cause of economic considerations, as,
parts) and also in the metal-working industry (fittings, gauge heads, accu-
at atmosphere
mulators).
r
X-ray formation
r
limited sheet thickness (max. 10 mm)
r
high investment
Under extreme demands on the weld-
r small working distance
ing time, reduced requirements to the weld geometry, distortion and in case
br-er9-13e.cdr
© ISF 2002
Disadvantages of EBW
of full material compatibility with air or shielding gas, out-of-vacuum welding Figure 9.13
9. Electron Beam Welding
123 units are applied, Figure 9.14. Their advantages are the continuous welding time and/or short cycle times. Areas of application are in the metalworking industry (precision tubes, bimetal strips) and in the automotive industry (converters, pinion cages, socket joints and module holders).
A further distinction criterion is the adjustment of the vacuum chambers to the different joining tasks. Universal machines are characterised by their simply designed working chamber, br-er9-14e_f.cdr
Figure 9.15. They are equipped with Machine Concept - Conventional Plant
vertically or horizontally positioned and, in most cases, travelling beam
Figure 9.14 generators. Here, several workpieces can be welded in subsequence during an evacuation cycle. The largest, presently existing working chamber has a volume of 265 m³.
Clock system machines, in contrast, are equipped with several small vacuum chambers which are adapted to the workpiece shape and they are, therefore,
characterised
by
short
evacuation times, Figure 9.16. Just immediately before the welding starts, is the beam gun coupled to the vac-
br-er9-15e_f.cdr
uum chamber which has been evacu-
EBW Clock System Machine
ated during the preceding evacuation Figure 9.15
9. Electron Beam Welding
124 cycle, while, at the same time, the next vacuum chamber may be flooded and charged/loaded.
Conveyor machines allow the continuous production of welded joints, as, for example, bimetal semi finished products such as saw blades or thermostatic bimetals, Figure 9.17. In the main chamber of these units is a gradually raising pressure system as partial vacsemi-finished material
endproduct
uum pre and post activated, to serve as a vacuum lock. Systems which are operating with a
br-er9-16e_f.cdr
mobile and local vacuum are characterEBW Conveyor Machine
Figure 9.16
ised by shorter evacuation times with a simultaneous maintenance of the vac-
uum by decreasing the pumping volume. In the “local vacuum systems”,
butt weld
with the use of suitable sealing, is the vacuum produced only in the welding area. In “mobile vacuum systems”
T-joint/ fillet weld
welding is carried out in a small vacuum chamber which is restricted to the welding area but is travelling along the
a)
b)
welded seam. In this case, a sufficient sealing between workpiece and vac-
T-joint butt welded
lap weld
uum chamber is more difficult. With these types of machine design, electron beam welding may be carried out with components which, due to
br-er9-17e.cdr
their sizes, can not be loaded into a
© ISF 2002
Seam Appearance for EB-Welding in Vacuum
stationary vacuum chamber (e.g. vesFigure 9.17
9. Electron Beam Welding
125 sel skins, components for particle accelerators and nuclear fusion plants).
In general the workpiece is moved during electron beam welding, while the beam remains stationary and is directed onto the workpiece in the horizontal or the vertical position. Depending on the control systems of the working table and similar to conventional welding are different welding positions possible. The weld type preferred in electron beam welding is the plain butt weld. Frequently, also cenbr-er9-18e.cdr
© ISF 2002
Seam Appearence at Atmospheric Welding (NV-EBW)
tring allowance for centralising tasks and machining is made. For the execution of axial welds, slightly over-
Figure 9.18 sized parts (press fit) should be selected during weld preparation, as a transverse shrinkage sets in at the 150
beginning of the weld and may lead to a considerable increase of the gap width in the opposite groove area. In some cases also
EBW MSG UP (narrow gap)(narrow gap)
T-welds may be carried out; the T-joint
EBW
with a plain butt weld should, however, be chosen only when the demands on the strength of the joints are low, Figure 9.18. As the beam spread is large
welding current
0,27 A
welding voltage groove area
150.000 V
UP (conventional)
MSG (narrow gap) 260 A
2
UP (narrow gap) 650 A
30 V
510 A
30 V 2
UP (conventional) 28 V
2
2
number of passes
896 mm 1
2098 mm 35
filler metal
0
melting efficiency
7,7 kg/h
energy input
64·10 kJ
128·10 kJ
293·10 kJ
377·10 kJ
welding time
27 min
4 h 35 min
4 h 11 min
7 h 20 min
3
4905 mm 81
5966 mm 143
23 kg
54 kg
66 kg
5 kg/h
13 kg/h 3
3
9 kg/h 3
under atmosphere, odd seam formations have to be considered during Non-Vacuum Electron Beam Welding,
br-er9-19e.cdr
© ISF 2002
Comparison of EB, GMAW and SAWNarrow Gap and Conventional SAW
Figure 9.19. Figure 9.19
9. Electron Beam Welding
126
In order to receive uniform and reproducible results with electron beam welding, an exact knowledge about the beam geometry is necessary and a prerequisite for:
- tests on the interactions between beam and substance - applicability of welding parameters to other welding machines - development of beam generation systems.
The objective of many tests is therefore the exact measurement of the beam and the investigation of the effects of different beam geometries on the welding result. For the exact measurement of the electron beam, a microprocessor-controlled measuring system has been developed in the ISF. The electron beam is linearly scanned at a high speed by means of a point probe, which, with a diameter of 20 µm is much smaller than the beam diameter in the focus, Figure 9.20. When the electron beam is deflected through the aperture diaphragm located inside the sensor, the electrons flowing through the diaphragm are picked up by a Faraday shield and industrial areas: l automotive industries l aircraft and space industries l mechanical engineering l tool construction l nuclear power industries l power plants l fine mechanics and electrical industries l job shop
diverted over a precision resistor. The time progression of the signal, intercepted at the resistor, corresponds with the intensity distribution of the electron beam in the scanning path. In order to receive an overall picture of
material: l almost all steels l aluminium and its alloys l magnesium alloys l copper and its alloys l titanium l tungsten l gold l material combinations (e.g. Cu-steel, bronze-steel) l ceramics (electrically conductive) br-er9-20e.cdr
the electron beam, the beam is line scanned over the slit sensor (60 lines). An evaluation program creates a perspective view of the power density distribution in the beam and also a two-dimensional representation of
EBW Fields of Application
Figure 9.20
the power density distribution inside
lines with the same power density.
9. Electron Beam Welding
127 An example for a measured electron
hole sensor hole with aperture diaphragm Faraday cup (20 µm)
beam is shown in Figure 9.21. It can track of the beam
cross section of the beam
measurement field
be seen clearly that the cathode had not
been
heated
up
sufficiently.
Therefore, the electrons are sucked off directly from the cathode surface during saturation and unsaturated beams, which may lead to impaired
slit sensor
welding results, develop. During the slit with Faraday cup
space charge mode of a generator,
cross section of the beam
the electron cloud is sufficiently large,
voltage
i.e., there are always enough electrons which can be sucked off. In the ideal case, the developed power denbeam deflection br-er9-21e.cdr
sity is rotationally symmetrical and in Two Principles of Electron Beam Measuring
accordance with the Gaussian distribution curve.
Figure 9.21 The electron signals are used for the automatic seam tracking. These may be either primary or secondary electrons or passing-through current or the developing X-rays. When backscattered primary electrons are used, the electron beam is scanned transversely to the groove. A computer may determine the position of the groove relative FILENAME: R I N G S T R Accel. voltage: 150 kV Beam current: 600 mA Prefocus current: 700 mA Main focus current: 1500 mA Cath. heat current: 500 mm Max. Density: 26,456 kW/mm2 2 Ref. Density: 26,456 kW/mm
to the beam by the signals from the reflected electrons. In correspondence with the deflection the beam is guided by electromagnetic deflection coils or by moving the working table.
br-er9-22e_f.cdr
© ISF 2002
Energy Concentration and Development in Electron Beam
This kind of seam tracking system may be used either on-line or off-line. Figure 9.22
9. Electron Beam Welding
128
The broad variation range of the weldable materials and also material thicknesses offer this joining method a large range of application, Figure 9.22. Besides the fine and micro welding carried out by the electronics industry where in particular the low heat input and the precisely programmable control is of importance, electron beam welding is also particularly suited for the joining of large cross-sections.
10. Laser Beam Welding
2003
10. Laser Beam Welding
129 The term laser is the abbreviation for
1917 postulate of stimulated emission by Einstein
,,Light Amplification by Stimulated
1950 work out of physical basics and realisation of a maser (Microwave Amplification by Stimulated Emission of Radiation) by Towens, Prokhorov, Basov
Emission of Radiation”. The laser is the further development of the maser
1954 construction of the first maser
(m=microwave),
1960 construction of the first ruby laser (Light Amplification by Stimulated Emission of Radiation)
Figure
10.1.
Al-
1961 manufacturing of the first HeNe lasers and Nd: glass lasers
though the principle of the stimulated
1962 development of the first semiconductor lasers
emission
1964 nobel price for Towens, Prokhorov and Basov for their works in the field of masers construction of the first Nd:YAG solid state lasers and CO2 gas lasers
mechanical fundamentals have al-
1966 established laser emission on organic dyes
the beginning of the 20th century, the
and
the
quantum-
ready been postulated by Einstein in
since increased application of CO2 and solid state laser 1970 technologies in industry
first laser - a ruby laser - was not
1975 first applications of laser beam cutting in sheet fabrication industry
implemented until 1960 in the Hughes
1983 introduction into the market of 1-kW-CO2 lasers
Research Laboratories. Until then
1984 first applications of laser beam welding in industrial serial production
numerous tests on materials had to
br-er10-01e.cdr
be carried out in order to gain a more History of the Laser
precise knowledge about the atomic structure. The following years had
Figure 10.1
been characterised by a fast devel-
opment of the laser technology. Already since the beginning of the Seventies and, increasingly since the Eighties when the first high-performance lasers were available, CO2 and solid state lasers have been used for production metal working. The number of the annual sales of la-
3
ser beam sources
10 €
has constantly in-
2
creased
the
1.5
course of the last
1
few
in
9
years,
Figure
0.5
10.2. 0
1986
1988
1990
1992
1994
Japan and South East Asia North America West Europe
The application arbr-er10-02e.cdr
eas for the laser beam sources sold Figure 10.2
1996
1998
2000
10. Laser Beam Welding
130
in 1994 are shown in Figure 10.3. The main application areas of the laser in the field of production metal working are joining and cutting jobs.
The availability of more efficient laser drilling 1,8%
welding 18,7%
beam
opens up new ap-
inscribe 20,5%
others 9,3%
sources
plication
possibili-
ties and - guided by
financial
con-
siderations - makes cutting 44,3%
micro electronics 5,4%
the use of the laser also more attrac-
br-er10-03e.cdr
tive, Figure 10.4. Figure 10.3
Figure 10.5 shows the characteristic properties of the laser beam. By reason of the induced or stimu40
lated emission the
kW
laser power
radiation is coher-
CO2
20 10
ent
5
chromatic. As the
4
divergence is only
3
1/10
Nd:YAG
2
mrad,
mono-
long
transmission paths
1 2000
1995
1990
1985
1980
1975
1970
diode laser 0
and
without
significant
beam divergences
br-er10-04e.cdr
are possible.
Figure 10.4
10. Laser Beam Welding
131
Inside the resonator, Figure 10.6, the laser-active medium (gas molecules, ions) is excited to a higher energy level (“pumping”) by energy input (electrical gas discharge, flash lamps).
During retreat to a lower level, the energy is released in the form of a light quantum (photon). The wave length depends on the energy difference between both excited states and is thus a characteristic for the respective laser-active medium. A distinction is made between spontaneous and induced transition. While the spontaneous emission is non-directional and in coherent (e.g. in fluorescent tubes) is a laser beam generated
by
induced
emission
when
a
particle
with
a
light bulb
Laser induced emission
E2
higher energy level
0,46"
is hit by a photon.
exited state
monochromatic
polychromatic
ton has the same
(multiple wave length)
coherent
incoherent
(fre-
(in phase)
(not in phase)
small divergence
large divergence
quency,
ground state
0,9 4"
The resulting pho-
properties
E1
direction, © ISF 2002
br-er10-05e.cdr
phase) as the excitCharacteristics of Laser Beams
ing photon (“coherence”). In order to
Figure 10.5
maintain the ratio of resonator
the desired induced
energy source
emission I spontaneous emission as as
possible,
the upper energy
laser beam
high
active laser medium
level must be constantly
over-
crowded, in comparison
with
the
fully reflecting mirror R = 100%
energy source
© ISF 2002
br-er10-06e.cdr
lower one, the socalled
Laser Principle
“laserFigure 10.6
partially reflecting mirror R < 100%
10. Laser Beam Welding
132
inversion”. As result, a stationary light wave is formed between the mirrors of the resonator (one of which is semi-reflecting) causing parts of the excited laser-active medium to emit light. In the field of production metal working, and particularly in welding, especially CO2 and Nd:YAG lasers are applied for their high power outputs. At present, the development of diode lasers is so far advanced that their sporadic use in the field of material processing is also possible. The industrial standard powers for CO2 lasers are, nowadays, approximately 5 - 20 kW, lasers with powers of up to 40 kW are available. In the field of solid state lasers average output powers of up to 4 kW are nowadays obtainable.
In the case of the 0,6
thrust of second type 2
energy
CO2 laser, Figure
002
eV transmission of vibration energy
0,4
transition without emission
10.7,
where
the
resonator is filled 0,3 0,288 eV
thrust of first type 1
001
∆E = 0,002 eV
0,290 eV
0,2
100
0,1
discharge through thrust with helium
0
with a N2-C02-He
LASER λ = 10,6 µm
0
gas mixture, pumping is carried out
000
N2
over the vibrational
CO2 © ISF 2002
br-er10-07e.cdr
Energy Diagram of CO2 Laser
excitation of nitrogen
molecules
which again, with
Figure 10.7
thrusts of the second type, transfer
radio frequency high voltage exitaion
their laser beam
vibrational
energy to the carbon dioxide. During
cooling water
cooling water
the transition to the laser gas
lower energy level, CO2
laser gas: CO2: 5 l/h He: 100 l/h N2 : 45 l/h
vakuum pump
gas circulation pump
emit
molecules a
radiation
with a wavelength
br-er10-08e.cdr
of 10.6 µm. The helium atoms, fiFigure 10.8
10. Laser Beam Welding
133 nally, lead the CO2 Cooling water
molecules back to
laser gas: CO2: 11 l/h He: 142 l/h N2: 130 l/h
turning mirrors
their energy level.
gas circulation pump laser beam
The efficiency of up mirror
to 15%, which is
(partially reflecting)
achievable
gas discharge
with
CO2 high performlaser gas
end mirror
br-er10-09e.cdr
ance lasers, is, in
cooling water
comparison
with
other
sys-
laser
tems, Figure 10.9
relatively
high. The high dissipation component
is the heat which must be discharged from the resonator. This is achieved by means of the constant gas mixture circulation and cooling by heat exchangers. In dependence of the type of gas transport, laser systems are classified into longitudinal-flow and transverse-flow laser systems, Figures 10.8 and 10.9.
With transverse-flow laser systems of a compact design can the multiple folding ability of the beam reach higher output powers than those achievable with longitudinal-flow systems, the beam quality, however, is worse. In d.c.-excited systems (high voltage), the f2,57"
electrodes are po-
d0
sitioned inside the unfocussed beam
ΘF
focussed beam
resonator. The indF
teraction
between
the electrode material and the gas 2λ 1 K= π d.σΘσ
molecules 0 90% - well suitable to automatic function
beam welding are depicted in Figure 10.25. Further relevant
- high power density - small beam diameter - high welding speed - non-contact tool - atmosphere welding possible
disadvantages
influential
- expensive edge preparation - exact positioning required - danger of increased hardness - danger of cracks - Al, Cu difficult to weld - expensive beam transmission and forming - power losses at optical devices - laser radiation protection required - high investment cost - low efficency (CO2-Laser: < 20%, Nd:YAG: < 5%) © ISF 2002
br-er10-24e.cdr
factors are, among
Advantages and Disadvantages of Laser Beam Welding
others, the material Figure 10.24
10. Laser Beam Welding
141 (thermal conductivity), the design of
28
penetration depth
the resonator (beam quality), the
0,2% C-steel CO2-laser
mm
(cross flow)
focal position and the applied optics
20
(focal length; focus diameter).
laser power:
16
15 kW
12 10 kW 8 kW 6 kW 4 kW
8 4
Figure 10.26 shows several joint
1,5 kW
0 0
0,6
1,2
1,8
m/min
shapes which are typical for car body
3,0
welding speed
production and which can be welded
penetration depth
15
by laser beam application.
X 5 CrNi 18 10 CO2-laser (axial flow)
mm
laser power:
5
0
6 kW
4 kW
2 kW 1 kW
0
1
2
3
4
5
6
7
m/min
9
welding speed
br-er10-25e.cdr
Penetration Depths
Figure 10.25 The high cooling rate during laser beam welding leads, when transforming steel materials are used, to significantly increased hardness values in comparison with other
welding
methods,
Figure
butt weld
fillet weld at overlap joint
10.27. These are a sign
for
the
in-
creased strength at a lower toughness and they are particularly
critical
circumstances
in
lap weld at overlap joint
of
dynamic loads. br-er10-26e.cdr
Figure 10.26
flanged weld at overlap joint
10. Laser Beam Welding
142
The small beam diameter demands the very precise manipulation and positioning of the workpiece or of 500 HV 0,4
the beam and an
WMA MAZ
MAZ
laser beam weld
hardness
exact weld preparation,
Figure
10.28.
Otherwise,
as result, lack of MAZ
MAZ
0
12
distance from the weld centre
submerged arc weld
weld
fusion,
sagged
welds or concave root surfaces are possible weld de-
submerged arc weld br-er10-27e_f.cdr
fects.
Figure 10.27
Caused by the high cooling rate and, in connection with this, the insufficient degassing of the molten metal, pore formation may occur during laser beam welding of, in particular, thick plates (very deep welds) or while carrying out welding-in works (insufficient degassing over the root), Figure 10.29. However, too low a weld speed may also cause pore formation when the molten metal picks up gases from the root side. The materials that misalignment
edge preparation
may
be
welded
with the laser reach from unalloyed and low-alloy steels up
(e ≤ 0,1 x plate thickness)
to high quality titagap
beam
mispositioning
nium
and
nickel
based alloys. The high carbon con(a ≤ 0,1 x plate thickness)
tent of the trans© ISF 2002
br-er10-28e.cdr
forming steel mateWelding Defects
Figure 10.28
rials is, due to the
10. Laser Beam Welding
143
high cooling rate, to be considered a critical influential factor where contents of C > 0.22% may be stipulated as the limiting reference value. Aluminium and copper properties cause problems during energy input and process stability. Highly reactive materials demand, also during laser beam welding, sufficient gas shielding beyond the solidification of the weld seam. The sole application of working gases is, as a rule, not adequate.
vw = 0,7 m/min
vw = 0,9 m/min
vw = 1,5 m/min
material: P460N (StE460), s = 20 mm, P = 15 kw © ISF 2002
br-er10-29e.cdr
Porosity
Figure 10.29
The application of laser beam welding may be extended by process variants. One is laser beam welding with filler wire, Figures 10.30 and 10.31 which offers the following advantages: - influence on the mechanic-technological properties of the weld and fusion zone (e.g. strength, toughness, corrosion, wear resistance) over the metallurgical composition of the filler wire - reduction of the demands on the accuracy of the weld preparation in regard to edge misalignment, edge preparation and beam misalignment, due to larger molten pools - “filling” of non-ideal, for example, V-shaped groove geometries - a realisation of a defined weld reinforcement on the beam entry and beam exit side.
10. Laser Beam Welding
144
The exact positioning of the filler wire is a prerequisite for a high weld quality and a sufficient dilution of the molten pool through which filler wire of different composition as the base can reach right to the root. Therefore, the use of sensor systems is indispensable for industrial application, Figure 10.32. The sensor systems are to take over the tasks of - process control, - weld quality as surance - beam positioning and joint tracking, respectively.
welding direction
filler wire
laser beam
filler wire
laser beam
gas
gas plasma
weld metal
molten pool
plasma
work piece
molten pool
keyhole
forward wire feeding
work piece
weld metal
keyhole
backward wire feeding
br-er10-30e.cdr
Figure 10.30
without filler wire
with filler wire
increase of gap bridging ability material: S380N (StE 380) gap: 0,5 mm PL = 8,3 kW VW = 3 m/min ES = 166 J/min s = 4 mm
filler wire: Sg2 dw = 0,8 mm
weld zone
Possibility of metallurgical influence
weld zone
material combination:
10CrMo9-10/ X6CrNiTi18-10 PL = 5,0 kW
br-er10-31e.cdr
Figure 10.31
gap: 0 mm vw = 1,6 m/min
gap: 0,5 mm wire: SG-Ni Cr21 Fe18 Mo
vw = 1,0 m/min dw = 1,2 mm
10. Laser Beam Welding
145
The present state-of-the-art is the further development of systems for industrial applications which until now have been tested in the laboratory. Welding by means of solid state lasers has, in the past, mainly been applied by manufacturers from the fields of precision mechanics and microelectronics. Ever since solid state lasers with higher powers are available on the market, they are applied in the car industry to an ever increasing degree. This is due to their more variable beam manipulation possibilities when comparing with CO2 lasers. The CO2 laser is mostly used by the car industry and with sensing device; fill factor 120 %
by
their
ancillary
KB 4620/9 20:1 10/92
industry for welding
KB 4620/6 20:1 10/92
KB 4620/4 20:1 10/92
KB 4620/0 20:1 10/92
KB 4620/41 20:1 10/92
KB 4620/38 20:1 10/92
Probe MS1-6C Probe MS1-5A Probe MS1-4C Probe MS1-3A Probe MS1-2B Probe MS1-1C
rotation-
0.1 mm
0.2 mm
0.3 mm
0.4 mm
0.5 mm
0.6 mm
KB 4620/12 20:1 10/92
KB 4620/17 20:1 10/92
KB 4621/15 20:1 10/92
KB 4621/12 20:1 10/92
KB 4621/9 20:1 10/92
KB 4621/7 20:1 10/92
symmetrical massproduced parts or sheets.
Figure
10.33 shows some typical
Probe OS1-6A Probe OS1-5C Probe OS1-4C Probe OS1-3B Probe OS1-2B Probe OS1-1B 1 mm
without sensing device; wire speed vD = 4 m/min constant
application
br-er10-32e.cdr
examples for laser beam welding. Figure 10.32
aerospace industry automotive industry
- engine components - instrument cases
- gear parts
steel industry - pipe production - vehicle superstructures - continuous metal strips - tins
(cog-wheels, planet gears)
- body-making (bottom plates, skins)
- engine components (tappet housings, diesel engine precombustion chambers)
electronic industry medical industry - heart pacemaker cases - artificial hip joints
plant and apparatus engineering
- PCBs - accumulator cases - transformer plates - CRTs
- seal welds at housings - measurement probes © ISF 2002
br-er10-33e.cdr
Practical Application Fields
Figure 10.33
11. Surfacing and Shape Welding
2003
11. Surfacing and Shape Welding
146 DIN 1910 (“Welding”) classifies the welding
base metal/ surfacing metal
according
process to
its
applications: weld-
similar
dissimilar
l for repair welding
l hardfacing (wear protection)
ing of joints and
l cladding (corrosion prevention)
surfacing. Accord-
l buffering (production of an appropriate-to-the-type-of-duty joint of dissimilar materials)
ing to DIN 1910 surfacing
is
the
coating of a workbr-er11-01e.cdr
piece by means of welding. Figure 11.1
Dependent on the
applied filler material a further classification may be made: deposition repair welding and surfacing for the production of a composite material with certain functions. Surfacing carried out with wear-resistant materials in preference to the base metal material is called hardfacing; but when mainly chemically stable filler materials are used, the method is called cladding. In the case of buffering, surfacing layers are produced which allow the appropriate-to-the-type-of-duty joining of dis-
m
wear caused by very high impact and compressive stress
similar materials and/or of materials
m
wear by friction (metal against metal) during high impact and compression stress
m
strong sanding or grinding wear
m
very strong wear caused by grinding during low impact stress
m
cold forming tools
m
hot forming tools
m
cavitation
strongly differing thermal expansion
m
wear parts (plastics industry)
coefficients.
m
corrosion
Figure 11.2 shows different kinds of
m
temperature stresses
with differing properties, Figure 11.1.
A buffering, for instance, is an intermediate layer made from a relatively tough material between two layers with
stresses which demand the surfacing of components. Furthermore surfacing
br-er11-02e.cdr
Components Kinds of Stress
may be used for primary forming as well as for joining by primary forming. Figure 11.2
11. Surfacing and Shape Welding
147 In case of surfacing - as for all fabrication processes - certain limiting conditions have to be observed. For ex-
component (material)
ample, hard and wear-resistant weld filler metals cannot be drawn into solid wires. Here, another form has to be
coating
stress compatibility
selected (filler wire, continuously cast manufacturing conditions availability
rods, powder). Process materials, as for example SA welding flux demand a certain welding position which in terms limits the method of welding.
coating material (filler)
consumable
surfacing method
The coating material must be selected with view to the type of duty and,
br-er11-03e.cdr
moreover, must be compatible with Boundary Conditions in Surfacing
the base metal, Figure 11.3.
Figure 11.3 For all surfacing tasks a large product line of welding filler metals is available. In dependence on the welding method as well as on the selected ma-
wearing protection (armouring) hard facing on
terials, filler metals in the form of wires,
q cobalt base
filler wires, strips, cored strips, rods or
q nickel base
powder are applied, Figure 11.4.
q iron base
The filler/base metal dilution is rather important, as the desired high-quality
corrosion prevention
properties of the surfacing layer dete-
q ferritic to martensitic chromium steel alloys
riorate with the increasing degree of
q soft martensitic chromium-nickel steel alloys
dilution.
q austenitic-ferritic chromium-nickel steel alloys q austenitic chromium-nickel steel alloys
A weld parameter optimisation has the br-er11-04e.cdr
objective to optimise the degree of dilu-
Materials for Surfacing
tion in order to guarantee a sufficient Figure 11.4
11. Surfacing and Shape Welding
148
adherence of the layer with the minimum metal dissimilation. A planimetric determination of the surfacing and penetration areas will roughly assess the proportion of filler to base metal. When the analysis surface built up by welding FB
of base and filler metal is known, a
penetration area FP
more precise calcubase metal
lation is possible by the determination of
FP FP + FB
AD=
the content of a cer-
(X-contentsurfacing layer - X-contentFM ) [% in weight] (X-contentbase metal - X-contentFM ) [% in weight]
AD =
tain element in the surfacing layer as
x 100%
FM: weld filler metal
x 100%
AD: dilution © ISF 2002
br-er11-05e.cdr
well as in the base Definition of Dilution
metal, Figure 11.5. Figure 11.5 Figure 11.6 shows record charts of an electron beam microprobe analysis for the elements nickel and chromium. It is Cr percentages by mass
30
evident that - after passing a narrow
%
transition zone between base metal
20
and layer — the analysis inside the layer is quasi constant.
10
0 0
100
200
300
µm
As depicted in Figure 11.7 almost all 500
arc welding methods are not only suit-
distance
Ni percentages by mass
30
able for joining but also for surfacing.
% 20
In the case of the strip-electrode submerged-arc surfacing process
10
normally strips (widths: 20 - 120mm) 0
are used. These strips allow high clad0
100
br-er11-06e.cdr
200
300
distance
Microprobe Analyses
µm
500
ding rates. Solid wire electrodes as well as flux-cored strip electrodes are used. The flux-cored strip electrodes contain
Figure 11.6
11. Surfacing and Shape Welding
149
certain alloying elements. The strip is continuously fed into the process via feed rollers. Current contact is normally carried out via copper contact jaws which in some cases are protected against
wear
by
metal-arc welding
hard metal inserts. The
- stick electrode - filler wire
slag-forming
arc welding with self-shielded cored wire electrode
flux is supplied onto the
workpiece
- filler wire
in
inert gas-shielded arc welding
submerged arc welding
- MIG / MAG - MIG cold wire - filler wire
- wire electrode - strip electrode
TIG welding
electroslag welding - wire electrode
- TIG cold wire
front of the strip arc spraying
plasma welding
electrode by means - powder - wire
of a flux support. The non-molten flux
- plasma powder - plasma hot wire plasma spraying
br-er11-07e.cdr
can
be
extracted
and returned to the flux circuit.
Figure 11.7
Should the slag developed on top of the welding bead not detach itself, it will have to be removed mechanically in order to avoid slag inclusions during overwelding. The arc wanders along the lower edge of the strip. Thus the strip is melted consistently, Figure 11.8.
power source drive rolls
+
-
filler metal
flux support
flux application slag surfacing bead base metal
br-er11-08e.cdr
Figure 11.8
11. Surfacing and Shape Welding
150
Figure 11.9 shows the cladding of a roll barrel. The coating is deposited helically while the workpiece is rotating. The weld head is moved axially over the workpiece.
br-er11-09e.cdr
Figure 11.9
The macro-section and possible weld defects of a strip-electrode submerged-arc surfacing process are depicted in Figure 11.10.
coarse grain zone
lack of fusion
mirco slag inclusions
sagged weld
base metal
crack formation in these areas of the coarse grain zone
br-er11-10e.cdr
Figure 11.10
gusset
undercuts
11. Surfacing and Shape Welding
151
Electroslag surfacing using a strip electrode is similar to strip-electrode SA surfacing, Figure 11.11. The difference is that the weld filler metal is not melted in the arc but in liquefied welding flux — the liquid slag – as a result of Joule resistance heating. The slag is held by a slight inclination
of
the
plate and the flux mound to prevent it from running off. molten pool
TIG weld surfacing is a suitable surfacing
method
br-er11-11e.cdr
for small and complicated Figure 11.11
contours
and/or low quantities
(e.g.
repair
work) with normally relatively low deposition rates. The process principle has already been shown when the TIG joint welding process was ex-
shielding gas nozzle
rod/ wire-shaped filler metal
plained, Figure 11.12. The arc is arc
burning between a gas-backed nonconsumable tungsten electrode and the workpiece. The arc melts the base metal and the wire or rod-shaped weld filler metal which is fed either continuously or intermittently. Thus a fusion welded joint develops between base metal and surfacing bead.
In the case of MIG/MAG surfacing
base metal (+ / ~)
tungsten electrode (- / ~) surfacing bead
br-er11-12e.cdr
© ISF 2002
Process Principle of TIG Weld Surfacing
processes the arc burns between a consumable wire electrode and the Figure 11.12
11. Surfacing and Shape Welding
152 workpiece.
This
method
allows
higher
contact tube
rates. Filler as well
wire feed device
shielding gas
deposition
shielding gas nozzle
as solid wires are
+ -
weld filler metal
power source
used.
arc
The
electrode
shielding gas
wire
has
a
positive, while the
surfacing bead
workpiece workpiece
oscillation
feed direction br-er11-13e.cdr
to
be
surfaced
has
a
negative
polarity,
Figure 11.13. Figure 11.13
A further development of the TIG welding process is plasma welding. While the TIG arc develops freely, the plasma welding arc is mechanically and thermally constricted by a water-cooled copper nozzle. Thus the arc obtains a higher energy density.
In the case of plasma arc powder surfacing this constricting nozzle has a positive, the tungsten electrode has a negative polarity, Figure 11.14. Through a pilot arc power supply a non-transferred arc (pilot arc) develops inside the torch. A second, separate power source feeds the transferred arc between electrode and workpiece. The non-transferred arc ionises the centrally
fed
plasma
gas
(inert
gases,
tungsten electrode
filler metal
plasma gas HIG
as, e.g., Ar or He)
UNTA
conveying gas power sources
thus
causing
a
shielding gas pilot arc welding arc
plasma jet of high energy to emerge from
the
This
plasma
surfacing bead
nozzle. jet
serves to produce
workpiece br-er11-14e.cdr
and to stabilise the Figure 11.14
oscillation
UTA
11. Surfacing and Shape Welding
153
arc striking ability of the transferred arc gap. The surfacing filler metal powder added by a feeding gas flow is melted in the plasma jet. The partly liquefied weld filler metal meets the by transferred arc molten base metal and forms the surfacing bead. A third gas flow, the shielding
gas,
protects
the surfacing bead and
the
section A
adjacent
ZW
high-temperature zone from the surrounding influence. The applied gases are
mainly
GW
inert
gases, as, for exbr-er11-15e.cdr
ample, Ar and He and/or Ar-/He mixFigure 11.15
tures.
The method is applied for surfacing small and medium-sized parts (car exhaust valves, extruder spirals). Figure 11.15 shows a cross-section of armour plating of a car exhaust valve seat. The fusion line, i.e., the region between surfacing and base metal, is shown enlarged on the right side of Figure 11.15 (blow-up). It shows hardfacing with cobalt which is high-temperature and hot gas corrosion resistant.
shielding gas
plasma power source
plasma gas
=
tungsten electrode
In plasma arc hot wire surfacing the base
metal
is
melted by an oscil-
arc wires from spool surfacing bead
lating plasma torch, Figure 11.16. The
~ workpiece
weld pool
hot wire power source
weld filler metal in the form of two parallel
wires
is
br-er11-16e.cdr
added to the base metal quite indeFigure 11.16
11. Surfacing and Shape Welding
154
pendently. The arc between the tips of the two parallel wires is generated through the application of a separate power source. The plasma arc with a length of approx. 20 mm is oscillating (oscillation width between 20 to 50 mm). The two wires are fed in a V-formation at an angle of approx. 30° and melt in the high-temperature region in the trailing zone of the plasma torch.
For surfacing purposes, besides the arc-welding methods, the beam welding methods laser beam and electron beam welding may also be applied. Figure 11.17 shows the process principle of laser surfacing. The powder filler metal is added to the laser beam via a powder nozzle and the powshielding gas nozzle
laser beam
powder nozzle
der gas flow is, in addition, constricted by
direction of the oscillation powder flow surfacing bead
shielding gas
shielding
gas
flow.
Friction
surfacing
is, in principle, simi-
Werkstück
lar to friction welding br-er11-17e.cdr
for the production of joints which due to the different materi-
Figure 11.17
als are difficult to produce with fusion electron beam
welding,
surface layer
11.18.
Figure
The filler metal is
base material
“advanced” over the metal foil
workpiece with high
metal foil feeding
pressure and rota-
workpiece
tion. By the pressure
feed direction
and ka11-18.cdr
Process Principle Electron Beam Surface Welding
Figure 11.18
the
relative
movement
frictional
© ISF 1998
heat develops and
11. Surfacing and Shape Welding
155
puts the weld filler end into a pasty condition. The advance motion causes an adherent, “spreaded” layer on the base metal. This method is not applied frequently and is mainly used for materials which show strong differences in their melting and oxidation behaviours.
A comparison of the different surfacing methods shows that the application fields are limited - dependent on the welding method. A specific method, for example, is the low filler/base metal dilution. These methods are applied where high-quality filler metals are welded. Another criterion for the selection of a surfacing method is the deposition rate. In the case of cladding large surfaces a method with a high deposition rate is chosen, this with regard to profitability.
In thermal spraying the filler metal is melted inside the torch and then, with a high kinetic energy, discharged onto the unfused but preheated workpiece surface.
There is no fusion of base and filler metal but rather adhesive binding and mechanical interlocking of the spray deposit with the base material. These mechanisms are effective only when the workpiece surface is coarse (pre-treatment by sandblasting) and free of oxides. The filler and base materials are metallic and non-metallic. Plastics may be sprayed as well. The utilisation of filler metals in thermal spraying is relatively low. The most important methods of thermal spraying are: plasma arc spraying, flame spraying and arc spraying. force filler metal
In powder flame
rotation advance
spraying an oxyacetylene flame provides the heating surfacing layer bulge
source where the centrally fed filler
base metal
metal
is
melted,
br-er11-18e.cdr
Figure 11.19. The kinetic energy for Figure 11.19
11. Surfacing and Shape Welding
156
the acceleration and atomisation of the filler metal is produced by compressed gas (air).
compressed air
spraying material
In contrast to pow-
workpiece
der flame spraying, is for flame spraying a wire filler metal fed mechanically into the centre cone, melted, at-
fuel gas-oxygen mixture
flame cone
omised and accel-
spray deposit
erated in direction
br-er11-19e.cdr
of
the
substrate,
Figure 11.20. Figure 11.20 In plasma arc spraying an internal, high-energy arc is ignited between the tungsten cathode and the anode, Figure 11.21. This arc ionises the plasma gas (argon, 50 100 l/min). The plasma emerges from the torch with a high kinetic and thermal energy and carries the side-fed powder along with it which then meets the workpiece surface in a semi-fluid state with the necessary kinetic energy. In the case of shape welding, steel shapes with larger dimensions and higher weights are produced from molten weld metal only. In comparison
compressed air
spraying jet
to cast parts this method
brings
about
essentially
more
favourable
non-binding sprayed particles (loss in spraying)
gas mixture
mechanotechnological mate-
adjustable wire feed device
rial properties, especially
a
spraying wire
better br-er11-20e.cdr
toughness
charac-
teristic. The reason Figure 11.21
fusing wire tip
spray deposit
11. Surfacing and Shape Welding
157
for this lies mainly in the high purity and the homogeneity of the steel which is helped by the repeated melting process and the resulting slag reactions. These properties are also put down to the favourable fine-grained structure formation which is achieved by the repeated subsequent thermal treatment with the multi-pass technique. Also in contrast with the shapes produced by forging, the workpieces produced by shape welding show quality advantages, especially in the isotropy and the regularity of their toughness and strength properties as far as larger workpiece thicknesses
are
con-
cerned. In Europe, powder injector
due to the lack of expensive
forging
equipment, high
back frame
isolation ring
gas middle distributor frame
anode carrier
very
copper anode
individual
weights may not be produced as forged jet of particles
parts. cooling water
Therefore,
shape
plasma gas
cooling water tungsten cathode
© ISF 2002
br-er11-21e.cdr
Plasma Powder Spraying Unit
welding is, for certain applications, a sensible logical nomical
Figure 11.22
technoand
eco-
alternative
primary forming (casting)
to the methods of primary
arc
forming,
forming or joining,
shape welding
Figure 11.22. forming (forging)
Figure 11.23 shows
joining (welding)
an early application which is related to
© ISF 2002
br-er11-22e.cdr
Shape Welding - Integration
the field of arts. Figure 11.23
11. Surfacing and Shape Welding
158
Baumkuchenmethode
+ several weld heads possible + no interruption during weld head failure - core made of foreign material necessary applications: shafts, large boiler shell rings, flanges
Töpfermethode
+ free rotationally-symmetrical shapes + several weld heads possible + weld head manipulation not necessary + each head capable to weld a specific layer + small diameters possible - component movement must correspond with the contour - number of weld heads limited when smaller diameters are welded applications: spherical caps, pipe bends, braces
Klammeraffe
+ transportable unit - limited welding efficiency applications: welding-on of connection pieces
br-er11-25e.cdr
br-er11-23e.cdr
© ISF 2002
Shape Welding Procedures
Shape Welded Goblet (1936)
Figure 11.24
Figure 11.25
The higher tooling costs in forging make the shape welding method less expensive; this applies to parts with certain increasing complexity. This comparison is, however, related to relatively low numbers of pieces, where the tooling costs per part are accordingly
higher,
Figure 11.24.
Figure 11.25 shows the principal procedure for the production
of
typical
shape-welded parts.
phase 7 phase 4 phase 2
joist
Cylindrical
containers are probr-er11-26e.cdr
duced
with
phase 5
phase 6
the
“BaumkuchenmeFigure 11.26
phase 3 traction mechanism
phase 1
turntable
11. Surfacing and Shape Welding
159 thode” method: the filler metal is welded by submerged-arc with helical
ti tes ng
1. welding of the half-torus 2. stress relief annealing 3. mechanical treatment 4. seperating/ halving 5. folding 6. welding togehter 7. stress relief annealing 8. testing
movement in multiple passes into a tube which has the function of a traction mechanism (for the most part mechanically removed later). This brings about the possibility to produce seamless containers with bottom and flange in one working cycle.
Elbows are mainly manufactured with the Töpfer method. On the traction mechanism a rotationally symmetrical part with a semicircle cross-section is br-er11-27e.cdr
produced which is later separated and Production of a Pipe Bend by Shape Welding
welded to an elbow, Figures 11.26 and 11.27. The Klammeraffe method
Figure 11.27
serves the purpose to weld external
connection pieces onto pipes. A portable unit which is connected with the pipe welds the connection pipe in a similar manner to the Töpfer method.
In the case of electron beam surfac-
forged products
€/kg
ing the filler metal is
added
to
the
process in the form
shape-welded products
pipe bends
braces
boiler shell rings
shafts
11.28.
spherical caps
of a film, Figure
complexity of the parts br-er11-24e.cdr
Figure 11.28
12. Thermal Cutting
2003
12. Thermal Cutting
160
Thermal cutting processes are applied in different fields of mechanical engineering, shipbuilding
and
process technology for the production Classification of thermal cutting processes - physics of the cutting process - degree of mechanisation
of components and for the preparation of welding edges. The thermal cut-
- type of energy source - arrangement of water bath
ting
processes
are classified into different categories
br-er12-01e.cdr
Classification of Thermal Cutting Processes acc. to DIN 2310-6
according to DIN 2310, Figure 12.1.
Figure 12.1 Figure 12.2 shows the classification according to the physics of the cutting process: - flame cutting – the material is mainly oxidised (burnt) - fusion cutting – the material is mainly fused - sublimation cutting – the material is mainly evaporat The gas jet and/or evaporation expansion is in all processes responsible for the ejection of molten material or emerging reaction products such as slag.
The different enFlame cutting The material is mainly oxidised;the products are blown out by an oxygen jet.
the thermal cut-
Fusion cutting The material is mainly fused and blown out by a high-speed gas jet.
ting are depicted in
Sublimation cutting The material is mainly evaporated. It is transported out of the cutting groove by the created expansion or by additional gas.
- gas,
Figure 12.3:
-
electrical
gas
discharge and - beams.
br-er12-02e.cdr
Classification of Processes by the Physics of Cutting
Figure 12.2
ergy carriers for
Electron
beams
for thermal cutting
12. Thermal Cutting
161
are listed in the DIN-Standard, they produce, however, only very small boreholes. Cutting is impossible.
Figure 12.4 depicts the different methods of thermal cutting with gas according to DIN 8580. These are: - flame cutting - metal powder flame cutting - metal powder
thermal cutting by:
fusion cutting
- gas - electrical gas discharge - sparks - arc - plasma
- flame planing -oxygen-lance cut-
- beams - laser beam (light) - electron beam - ion beam
ting - flame gouging or scarfing br-er12-03e.cdr
-flame cleaning
Classification of Thermal Cutting Processes acc. to DIN 2310-6
Figure 1.3 In flame cutting (principle is depicted in Figure 12.5) the material is brought to the ignition temperature by a heating flame and is then burnt in the oxygen stream. During the process the ignition temperature is maintained on the plate top side by the heating flame and below the plate top thermal cutting processes using gas:
side
by
thermal
conduction
and l
convection.
oxygen cutting
l
metal powder
l
flame cutting
fusion cutting
metal powder
However, this process
is
suited
for
l
flame planing
automation and is, also easy to apply
l
oxygen-lance cutting
l
flame gouging
l
l
flame cleaning
scarfing
br-er12-04e.cdr
on site. Figure 12.6.
Thermal Cutting Processes Using Gas
shows a commerFigure 1.4
12. Thermal Cutting
162
cial torch which combines a welding with a cutting torch. By means of different nozzle shapes the process may be adapted to varying materials and plate thicknesses. Hand-held or
torches cutting oxygen heating oxygen gas fuel
machine-type
torches
are
equipped with difheating flame
ferent cutting nozzles: Standard or block-type nozzles (cutting-oxygen pressure 5 bar) are
cutting jet
used for hand-held torches
and
workpiece
br-er12-05e.cdr
for
Principle of Oxygen Cutting
torches which are fixed to guide car-
Figure 12.5
riages.
The high-speed cutting nozzle (cutting-oxygen pressure 8 bar) allows higher cutting speeds with increased cutting-oxygen pressure. The heavy-duty cutting nozzle (cutting-oxygen pressure 11 bar) is mainly applied for economic cutting with flamecutting machines. A further development of the heavy-duty nozzle is the oxygenshrouded nozzle which allows even faster and more economic cutting of plates within
certain
thickness cutting oxygen
ranges.
Gas mixing is ei-
heating oxygen
ther carried out in
gas fuel mixing chamber
the torch handle, the cutting attachment,
manual cutting equipment as a cutting and welding torch combination
the
torch
head or in the nozzle gas mixing nozzle
block-type nozze
(gas
mixing
nozzle); in special
br-er12-06e.cdr
Cutting Torch and Nozzle Shapes
cases also outside the torch – in front
Figure 12.6
12. Thermal Cutting
163
of the nozzle. As the design of cutting torches is not yet subject to standardisation, many types and systems exist on the market.
The selection of a heating and cutting nozzle
torch
nozzle-to-work distance
torch kerf width
cutting jet
kerf
or
nozzles
important and depends mainly on
start
the cutting thick-
cut thickness
ness, the desired
cutting le
cut lengt h ngth
cutting
quality,
and/or the geometry of the cutting
end of the cut
br-er12-07e.cdr
Flame Cutting Terms
edge. Figure 12.7 gives a survey of the definitions of
Figure 12.7
flame-cutting.
In flame cutting, the thermal conductivity of the material must be low enough to constantly maintain the ignition temperature, Figure 12.8. Moreover, the material must neither melt during the oxidation nor form high-melting oxides, as these would produce difficult cutting surfaces. In accordance, only steel or titanium materials fulfill the conditions for oxygen cutting., Figure 12.9
The heating flame has to perform the following tasks:
Steel materials with a C-content of up
- rapid heating of the material (about 1200°C) - substitution of losses due to heat conduction in order to maintain a positive heat balance
to approx. 0.45% may be flame-cut
- preheating of cutting oxygen
without preheating,
- stabilisation of the cutting oxygen jet; formation
with a C-content of
of a cylindrical geometry over a extensive length and protection against nitrogen of the surrounding air
approx.
1.6%
flame-cutting carried
br-er12-08e.cdr
Function of the Flame During Flame Cutting
out
preheating,
is with be-
cause an increased Figure 12.8
12. Thermal Cutting
164
C-content demands more heat. Carbon accumulates at the cutting surface, so a very high degree of hardness is to be expected. Should the carbon content exceed 0.45% and should the material not have been subject to prior heat treatment, hardening cracks on the cutting
surface
are
regarded as likely.
The material has to fulfill the following requirements:
Some
- the ignition temperature has to be lower than the
alloying
elements
melting temperature - the melting temperature of the oxides has to be lower
high-melting
than the melting temperature of the material itself
form ox-
- the ignition temperature has to be permanently maintained;
ides which impair
i. e. the sum of the supplied energy and heat losses due to
the slag expulsion
heat conduction has to result in a positive heat balance
and influence the thermal conductiv-
br-er12-09e.cdr
ity.
Conditions of Flame Cutting
Figure 12.9
The iron-carbon equilibrium diagram illustrates the carbon content-temperature interrelation, Figure 12.10. As the carbon content increases, the melting temperature is lowered. That means: from a certain carbon content upwards, the ignition temperature is higher than the melting temperature, i.e., this would be the first violation to the basic requirement in flame cutting.
Steel compositions
steel
temperature [°C]
may influence flame cuttability
substan-
tially - the individual alloying
cast iron
1500
elements
may show recipro1000
liquid pasty
solid
Liquidus
rve n cu o i t i ign
Solidus solid
cate effects (reinforcing/weakening), 2,0
Figure 12.11. The
carbon content [%]
br-er12-10e.cdr
content limits of the alloying
Ignition Temperature in the Iron-Carbon-Equilibrium Diagram
constituFigure 12.10
12. Thermal Cutting
165
ents are therefore only reference values for the evalua-
Maximum allowable contents of alloy-elements:
tion of the flame
carbon:
cuttability of steels,
silicon:
up to 2,5 % with max. 0,2 %C
manganese:
up to 13 % and 1,3 % C
chromium:
up to 1,5 %
tungsten:
up to 10 % and 5 % Cr, 0,2 % Ni, 0,8 % C
nickel:
up to 7,0 % and/or up to 35 % with min. 0,3 % C
deteriorating, as a
copper:
up to 0,7 %
rule
molybdenum: up to 0,8 %, with higher proportions of W, Cr and C
as the cutting quality is substantially
already
with
alloy
con-
up to 1,6 %
not suitable for cutting
lower
br-er12-11e.cdr
tents.
Flame Cutting Suitability in Dependance of Alloy-Elements
Figure 12.11
By an arrangement of one or several
nozzles already during the cutting phase a weld preparation may be carried out and certain welding grooves be produced. Figure 12.12 shows torch arrangements for - the square butt weld, - the single V butt weld, - the single V butt weld with root face, - the double V butt weld and -
the double V butt weld with root face.
It has to be considered that, particularly in cases where flame cutting is applied
for
weld
square butt weld
single-V butt weld
single-V butt weld with rootface
preparations, flame cutting-related
de-
fects may lead to increased
weld
dressing
work.
double-V butt weld
double-V butt weld with root face
br-er12-12e.cdr
Slag adhesion or
Weld-Groove Preparation by Oxygen Cutting
chains of molten Figure 12.12
12. Thermal Cutting
166 globules have to be removed in
cratering: sporadic craterings connected craterings cratering areas
edge defect: edge rounding chain of fused globules edge overhang
order
to
guarantee process safety
adherent slag: slag adhearing to bottom cut edge cut face defects: kerf constriction or extension angular deviation step at lower edge of the cut excessive depth of cutting grooves
and part accuracy
cracks: face cracks cracks below the cut face
for
the
subsequent processes. Figure
br-er12-13e.cdr
12.13
gives a survey
Possible Flame Cutting Defects
of Figure 12.13
possible
defects
in
flame cutting.
In order to improve the flame-cutting capacity and/or cutting of materials which are normally not to be flame-cut the powder flame cutting process may be applied. Here, in addition to the cutting oxygen, iron powder is blown into the cutting gap. In the flame, the iron powder oxidises very fast and adds further energy to the process. Through the additional energy input the
high-melting
oxides of the highalloy materials are molten.
oxygen water seperator
compressed air
acetylene
Figure
12.14 shows a diagrammatic
powder dispenser
repre-
sentation of a metal powder
cutting
br-er12-14e.cdr
arrangement.
Metal Powder Flame Cutting
Figure 12.14
12. Thermal Cutting
167
Figure 12.15 shows the
principle
of flame gouging
flame gouging and scarfing.
Both
scarfing
gas-heat oxygen mixture
methods are suited
gas-heat oxygen mixture
gouging oxygen
for the weld prepa-
scarfing oxygen
ration; material is removed
but
not
cut. This way, root passes
may
be br-er12-15e.cdr
grooved out or fil-
Flame Gouging and Scarfing
lets for welding may be produced later.
Figure 12.15
Figure 12.16 shows the methods of thermal cutting processes by electrical gas discharge: -
plasma cutting with non-transferred arc
-
plasma cutting with transferred arc
-
plasma cutting with transferred arc and secondary gas flow
-
plasma cutting with transferred arc and water injection
-
arc air gouging (represented diagrammatically)
-
arc oxygen cutting (represented diagrammatically) In plasma cutting the entire workpiece
Thermal cutting processes by electrical gas discharge:
must be heated to plasma cutting
- with non-transferred arc - with transferred arc -with secondary gas flow -with water injection
arc air gouging
arc oxygen cutting
the melting temperature by the plasma
carbon electrode compressed air
cutting oxygen
=
electrode coating
jet. The nozzle forms the plasma jet only
tube arc
in a restricted way and limits thus the cutting
ability
of
br-er2-16e.cdr
Thermal Cutting Processes by Electrical Gas Discharge
Figure 12.16
plate to a thickness of approx. 150 mm,
12. Thermal Cutting
168
Figure 12.17. Characteristic for the plasma cut are the cone-shaped formation of the kerf and the rounded edges in the plasma jet entry zone which were caused by the hot gas shield that envelops the plasma jet. These process-specific disadvantages may be significantly reduced or limited to just one side of the plate (high quality or scrap side), respectively, by the inclination of the torch and/or water addition. With the plasma cutting process, all electrically
conductive
materials may be separated.
Non-
conductive
materi-
als, or similar mate-
plasma gas
electrode
-
cooling water
power source
HF R
+
nozzle
rials, may be separated by the emergworkpiece
ing plasma flame, br-er12-17e.cdr
but only with limited
Plasma Cutting
ability. Figure 12.17 In order to cool and to reduce the emissions, plasma torches may be surrounded by additional gas or water curtains which also serve as arc constriction, Figure 12.18. In dry plasma cutting where Ar/H2, N2, or air are used, harmful substances always develop which not plasma gas
electrode
only have to be sucked
off
very
carefully but which water curtain
also must be discutting water swirl chamber
nozzle
cone of water
posed of. In
water-induced
plasma water bath workpiece
cutting
(plasma arc cutting in water or under
br-er12-18e.cdr
Water Injection Plasma Cutting
water) gases, dust, also the noise, and
Figure 12.18
12. Thermal Cutting
169
the UV radiation are, for the most part, held back by the water. A further, positive effect is the cooling of the cutting surface, Figure 12.18. Careful disposal of the residues
is
here
cutting with water bath
water injection plasma cutting with water curtain
plasma cutting with workpiece on water surface
underwater plasma cutting
inevitable.
Figure 12.19 gives a survey of the different cutting methods using a water
br-er12-19e.cdr
Types of Water Bath Plasma Cutting
bath. Figure 12.19
Figure 12.20 shows a torch which is equipped with an additional gas supply, the socalled secondary gas. The secondary gas shields the plasma jet and increases the transition resistance at the nozzle front. The so-called “double and/or parasite arcs” are avoided and nozzle life is increased. Thanks to new electrode materials, compressed air and even pure oxygen may be applied as plasma gas – therefore, in flame cutting, the burning of unalloyed steel may be used for increased capacity and
quality.
plasma gas
The
selection
of
plasma
forming
electrode
the
gases depends on
secondary gas
the requirements of
nozzle
the cutting process. Plasma
forming workpiece
media are argon, br-er12-20e.cdr
helium, hydrogen, Plasma Cutting With Secondary Gas Flow
nitrogen, air, oxygen or water.
Figure 12.20
12. Thermal Cutting
170
The advantage of the use of oxygen as plasma gas is in the achievable cutting speeds within the plate thickness range of approx. 3 – 12 mm (400 A, WIPC). In the steel plate thickness range of approx. 1 – 10 mm the application of 40 A-compressed air units is recomIn
com-
parison with 400 A WIPC
systems,
these allow vertical and
significantly
narrower
cutting
cutting speed [m/min]
mended.
machine type and plasma medium 1 WIPC, 400 A, O2 2 WIPC, 400 A, N2 3 200 A, s < 8 mm: N2 s > 8 mm: Ar/H2 4 40 A, compressed air
8 1 6 2 4 2
3 4
kerfs, but with lower 5
cutting speeds. Figure 12.21
shows
different
cutting
plate thickness [mm]
Cutting Speeds of Different Plasma Cutting Equipment for Steel Plates
Figure 12.21
gases.
In the thermal cutting with
Thermal cutting processes by laser beam
processes beams
only - laser beam combustion cutting
the laser is used as the jet generator for cutting,
- laser beam fusion cutting
Figure - laser beam sublimation cutting
12.22. Variations
of
the
15
br-er12-21e.cdr
speeds for different units and plasma
10
br-er12-22e.cdr
laser beam cutting
Thermal Cutting With Beams
process: Figure 12.22
-
laser beam combustion cutting, Figure 12.25
-
laser beam fusion cutting, Figure 12.26
-
laser beam sublimation cutting, Figure 12.27.
20
12. Thermal Cutting
171
The process sequence in laser beam combustion cutting is comparable to oxygen cutting. The material is heated to the ignition temperature and subsequently burnt in the oxygen stream, Figure 12.23. Due to the concentrated energy input almost all metals in the plate thickness range of up to approx. 2 mm may be cut. In addition, it is possible to achieve very good bur-free cutting qualities for stainless steels (thickness of up to approx. 8 mm) and for structural steels (thickness of up to 12 mm). Very narrow and parallel cutting kerfs are characteristic for laser beam cutting of structural steels.
In laser beam cutlens
ting, either oxygen (additional
energy
contribution for oxi-
cutting oxygen
dising materials) or an inactive cutting gas may be applied
laser focus thin layer of cristallised molten metal
workpiece
depending on the slag jet
cutting job. Besides, br-er12-23e.cdr
the very high beam Laser Beam Cutting
powers (pulsed/superpulse
Figure 12.23
d mode of operation) allow a direct evaporation of the
80
ting
and
cutlaser
beam sublimation
20
cutting the reflexion
of
the
40
laser
evaporating
combustion
60
melting
tion). In laser beam
heating-up
(sublimaabsorption factor
material
r) G-lase (Nd:YA 6 µm er) s a λ = 1,0 -l (CO 2 ,06 µm λ = 10
melting point Tm
boiling point Tb
temperature
beam of more than
br-er12-24e.cdr
Qualitative Temperature Dependency on Absorption Ability
90 % on the workpiece surface deFigure 12.24
12. Thermal Cutting
172
creases unevenly when the process starts. In laser beam fusion cutting remains the reflexion on the molten material, however, at more than 90%! Figure 12.24 shows the absorption factor of the laser light in dependence on the temperature. This factor mainly depends on the wave length of laser cutting (with oxygen jet) - the laser beam is focused on the workpiece surface and the material burns in the oxygen jet starting from the heated surface materials: - steel aluminium alloys, titanium alloys
the
used
light.
laser
When
the
melting point of the material has been reached, the ab-
cutting gas: - O2, N2, Ar criteria: - high cutting speed, cut faces with oxide skin br-er12-25e.cdr
Characteristics of the Laser Beam Cutting Processes I
sorption
factor
increases
un-
evenly
and
reaches values of more than 80%.
Figure 12.25
During laser beam combustion cutting of structural steel high cutting speeds are achieved due to the exothermal energy input and the low laser beam powers, Figure 12.25. In the above-mentioned case (dependent on beam quality, focussing, etc.), above a beam power of approx. 3,3 kW, spontaneous evaporation of the material takes place and allows sublimation cutting. Significantly higher laser powers are necessary to fuse the laser fusion cutting: - the laser beam melts the entire plate thickness (optimum focus point 1/3 below plate surface) - high reflection losses (>90%) materials: - metals, glasses, polymers
material and blow it out with an inert gas, as the reflexion loss remains constant.
cutting gas: - N2, Ar, He criterions: - cutting speed is only 10-15% in comparison to cutting with oxygen jet, characteristics melting drag lines
Important ence for
influ-
quantities the
cutting
br-er12-26e.cdr
Characteristics of the Laser Beam Cutting Processes II
Figure 12.26
speed and quality in laser beam cut-
12. Thermal Cutting
173
ting are the focus intensity, the position of the focus point in relation to the plate surface and the formation of the cutting gas flow. A prerequisite for a high intensity in the focus is the high beam quality (Gaussian intensity distribution in the beam) with a high beam power and suitable focussing optics. Laser beam cutting of contours, especially of pointed corners and narrow root faces, requires adaptation of the beam power in order to avoid heat accumulation and burning of the material. In such a case the beam power might be reduced in the continuous wave (CW) operating mode. With a decreasing beam efficiency decreases the cuttable plate thickness as well. Better suited is the switching of the laser to pulse mode (standard equipment of
laser evaporation cutting: - spontaneous evaporation of the material starting from 105 W/cm2 with high absorption rate and deep-penetration effect - metallic vapour is pressed from the cavity by own vapour pressure and by a supporting gas flow materials: - metals, wood, paper, ceramic, polymer
HF-excited lasers) where pulse height can right
be
selected
up
to
the
height of the con-
cutting gas: - N2, Ar, He (lens protection)
tinuous criteria: - low cutting speed, smooth cut edges, minimum heat input
wave.
super
pulse
equipment
br-er12-27e.cdr
A
(in-
creased excitation)
Characteristics of the Laser Beam Cutting Processes III
allows significantly higher pulse effi-
Figure 12.27
ciencies to be selected than those
laser 600 W 1500 W 600 W 1500 W 1500 W
steel
achieved with CW.
Cr-Ni-steel
Further
aluminium
of
application for the
plasma 50 A 5 kW 250 A 25 kW 500 A 150 kW
steel Cr-Ni-steel aluminium
pulse pulse
Stahl Cr-NiStahl
oxy-flame
1
and
super
operation
mode are punching 10
100 plate thickness [mm]
br-er12-28e.cdr
Fields of Application of Cutting Processes
Figure 12.28
fields
1000
and
laser
beam
sublimation cutting.
12. Thermal Cutting
174
Laser beam cutting of aluminium plates thicker than appx. 2 mm does not produce bur-free results due to a high reflexion property, high heat conductivity and large temperature
dif-
ferences between
CO2-laser (1500 W)
Al and Al2O3. The of
cuttig speeds [m/min]
addition
10
iron
powder allows the flame
cutting
stainless
of
steels
plasma cutting (WIPC, 300-600 A)
1 oxygen cutting (Vadura 1210-A)
(energy input and 0,1
improvement of the molten-metal
10
1
vis-
100
plate thickness [mm] br-er12-29e.cdr
cosity). The cutting quality,
Cutting Speeds of Thermal Cutting Processes
however,
does not meet high
Figure 12.29
standards.
Figure 12.28 shows a comparison of the different plate thicknesses which were cut using different processes. For the plate thickness range of up to 12 mm (steel plate), laser beam cutting is the approved precision cutting process. Plasma cutting of plates > 3 mm allows higher cutting speeds, in comparison to laser beam cutting, the cutting quality, however, is costs [DM/m cut length]
significantly
total costs
6
lower. Flame cut-
machine costs
5
ting is used for
4
cutting
laser
3
flame cutting with 3 torches
plasma
2
> 3 mm, the cutting speeds are,
1
in comparison to 5
10
15
20
25
30
35
40
plasma
plate thickness [mm] br-er12-30e.cdr
cutting,
significantly lower.
Thermal Cutting Costs - Steal
Figure 12.30
plates
With
increasing
an plate
thickness the dif-
12. Thermal Cutting
175
ference in the cutting speed is reduced. Plates with a thickness of more than 40 mm may be cut even faster using the flame cutting process.
Figure 12.29 shows the cutting speeds of some thermal cutting processes.
Apart from technological aspects, financial considerations as well determine the application of a certain cutting method. Figures 12.30 and 12.31 show a comparison of the costs of flame cutting, plasma arc and laser beam cutting – the costs per m/cutting extract from a costing acc. to VDI 3258
length
and the costs per
flame cutting (6-8 torches)
plasma cutting (plasma 300A)
laser beam cutting (laser 1500W)
170,000.00
220,000.00
500,000.00
investment total (replacement value)
€
calculation for a 6-yearaccounting depreciation
€/h
23.50
29.00
65.00
maintenance costs
€/h
3.50
4.00
10.00
energy costs
€/h
1.00
2.50
2.50
production cost unit rate costs/1 operating hour
€/h
65.00
75.00
130.00
operating The
hour.
high
invest-
ment costs for a laser beam cutting equipment
might
be a deterrent to 1 shift, 1600h/year, 80% availability, utilisation time 1280h/year
exploit cutting
the
high
qualities
br-er12-31e.cdr
Cost Comparison of Cutting Processes
obtainable with this process.
Figure 12.31
13. Special Processes
2003
13. Special Processes
175
Apart from the welding processes explained earlier there is also a multitude of special welding processes. One of them is stud welding. Figure 13.1 depicts different stud shapes. Depending on the application, the
studs
are
equipped with either internal or external
screw
threads; also studs with pointed tips or with
corrugated
shanks are used.
Figure 13.1 In arc stud welding, a distinction is basically made between three process variations. Figure 13.2. depicts the three variations – the differences lie in the kind of arc ignition and in the cycle of motions during the welding process.
Figure 13.2
13. Special Processes
176
The switching arrangement of an arc stud welding unit is shown in Figure 13.3. Besides a power source which produces high currents for a short-time, a control as well as a lifting device are necessary.
Figure 13.3
In drawn-arc stud welding the stud is first mounted onto the plate, Figure 13.4. The arc is ignited by lifting the stud and melts the entire stud diameter in a short time. When
stud
and
base
plate
are
fused, the stud is dipped
into
the
molten weld pool while the ceramic ferrule is forming the weld. After the solidification of the liquid weld pool the ceramic ferrule is knocked off. Figure 13.4
13. Special Processes
177
Figure 13.5 illustrates tip ignition stud welding. The tip melts away immediately after touching the plate and allows the arc to be ignited. The lifting of the stud is dispensed with. When the stud base is molten, the stud is positioned onto the partly molten workpiece.
Studs with diameters of up to 22 mm can be used. Welding currents of more than 1000 A are necessary.
The arc stud welding process allows to join
different
materials,
see
Fig-
ure 13.6. Problematic are the different melting points and the heat dissipation of the individual materials. Aluminium studs, for example, may not be welded onto steel.
The relatively high welding currents in the arc stud welding process cause the
somewhat
troublesome
Figure 13.5
sideeffects of the arc blow. Figure 13.7 depicts
different
arrangements current
of
contact
points and cable runs and illustrates the developing arc deflection (B,C,E). A, D and F show possible measures. Figure 13.6
counter-
13. Special Processes
178
In high-frequency welding of pipes the energy input into the workpiece may be carried out via sliding contacts, as shown in Figure 13.8, or via rollers, as shown in Figure 13.9. Only the high-frequency technique allows a safe current transfer in spite of the scale or oxide
layers.
Through the skin effect the current flows only conditionally at the surface. Therefore no thorough fusion of thick-wall
pipes
may be achieved.
Figure 13.7
Figure 13.8
Figure 13.9
13. Special Processes
179
Only welding of small wall thicknesses is profitable – as the weld speed must be greatly reduced with increasing wall thicknesses, Figure 13.10.
In induction welding – a process which is used frequently nowadays – the energy input is received contactless, Figure 13.11. Varying magnetic fields produce eddy currents inside the workpiece, which again cause resistance heating in the slotted tube. A distinction is made between coil inductors (left) and line inductors (right).
Figure 13.10
Also in case of induction welding flows the current flows only close to the surface areas of the pipe. Only the current part which reaches the joining zone and causes to fill the gap may be utilised.
Fig-
ure 13.12
illus-
trates two current paths. On the left side:
the
current
useful
path,
on
the right side: the useless
current
path which does not contribute to the fusion of the Figure 13.11
edges.
13. Special Processes
180
Figure 13.13 shows the effective depth during the inductive heating for different materials,
in
de-
pendence
on
the
frequency. As soon as the Curie temperature
point
is
reached, the effective depth for ferritic steels increases. Figure 13.12
Figure 13.13
Figure 13.14
13. Special Processes
181 The application of the induction welding method allows high
welding
speeds
of
than
more
100m/min,
Figure 13.14.
Aluminothermic fusion welding or cast
welding
mainly Figure 13.15
used
joining
is for
railway
tracks on site. A crucible is filled with a mixture consisting of aluminium powder and iron oxide. An exothermal reaction is initiated by an igniter – the aluminium oxidises and the iron oxide is reduced to iron, Figure 13.15. The molten iron flows into a ceramic mould which matches the contour track.
of
the
After
the
melt has cooled, the
mould
is
knocked off. Figure 13.16 the
process
sembly.
Figure 13.16
shows as-
13. Special Processes
182
Explosion welding or explosion cladding
is
fre-
quently used for joining dissimilar materials, as, for example,
unal-
loyed steel/alloyed steel,
cop-
per/aluminium
or
steel/aluminium. The
materials
which are to be
Figure 13.17
joined are pressed together
by
a
shock
wave.
Wavy
transitions
develop
in
the
joining area, Figures 13.17
and
13.18.
Figure 13.18
The determined cladding speed must be strictly adhered to during the welding process. If the welding speed is too low, lack of fusion is the result. If the welding speed is exceeded, the development of the waves in the joining zone is erratic. Figure 13.19 shows the critical cladding speeds for different material combinations.
13. Special Processes
Figure 13.19
183
Figure 13.20
Figure 13.20 shows a diagrammatic representation of a diffusion welding unit. Diffusion welding, like ultrasonic welding, is welding in the solid state. The surfaces which are to be joined are cleaned, polished and then joined in a vacuum with pressure and temperature. After a certain time (minutes, right up to several days) joining is achieved by diffusion processes.
The advantage of this costly welding method lies in the possibility of joining dissimilar materials without taking the risk of structural transformation due to the Figure 13.21
13. Special Processes heat
input.
ure 13.21
184
Figshows
several
possible
material
combina-
tions. The joining of two extremely different materials, as, e.g. austenite and a zirconium
alloy,
may be obtained by several
intermedi-
ate layers. Figure 13.22
Figure 13.22 shows the structure of a joint where nickel, copper and vanadium had been used as intermediate layers. As the diffusion of the individual components takes place only in the region close to the surface, very thin layers may be realised.
In cold pressure welding - in contrast to diffusion welding - a deformation is produced by the high contact pressure in the bonding plane, Figure 13.23. The joint surfaces
are
moved very close towards
each
other, i.e., to the atomic
distance.
Through transposition processes as well
as
through
adhesion
forces
can joining of similar and dissimilar materials be realFigure 13.23
ised.
13. Special Processes
185
Ultrasonic welding is used as a microwelding method. The process principle is shown in Figure 13.24. The surface layers of overlap arranged plates are destroyed by applying mechanical vibrator energy. At this instance are joining surfaces deformed by very short localised
warming
up
and
point-
interspersed connected. The joining members are welded under pressure, where one part small amplitudes (up to 50 µm) relative to the other is moved with with ultrasonic frequency. As far as metals are concerned, the vibratory vector is
in the joining zone, in contrast to ultrasonic welding of plastics. The ultrasonics which have been produced by a magnetostrictive transducer and transmitted by a sonotrode lie in the Figure 13.24
frequency range of 20 up to 60 Hz.
Figure 13.25 shows possible
material
combinations ultrasonic
for
weld-
ing. Further microwelding processes are methods which are also called heated element
welding
methods,
as,
for
example,
nailhead
bonding and wedge
Figure 13.25
13. Special Processes
186
bonding. These methods are applied in the electronics industry for joining very fine wires, as, for example, gold wires from microchips with aluminium strip conductors.
In wedge bonding a wire is positioned onto
the
contact
point via a feeding nozzle. The welding wedge is lowered and
the
welded
wire with
is the
aluminium thin foil, Figure 13.26.
The
wire is cut with a cutting tool.
Figure 13.26
In nailhead bonding, the wire which emerges from the feeding nozzle may have diameters from 12 to 100 µm. By a reducing hydrogen flame its end is molten to a globule, Figure 13.27. The nozzle then presses this globule onto the part aimed at and shapes it into a nail head.
Figure 13.28
de-
picts this type of weld.
A further method related to welding is soldering. The process of
principle
soldering
is
briefly explained in Figure 13.29. Figure 13.27
13. Special Processes
187
The individual soldering methods are classified into different mechanisms depending on the type of heating, Figure 13.30. There are two basic distinctions: soft soldering (melting temperature of the solder is approx. up to 450°C) and brazing (melting temperature of the brazing solder is approx.
up
to
1100°C. For hightemperature
sol-
dering solders with high melting points (melting
tempera-
ture is approx. up to 1200°C) are used. This process is frequently subject to automation.
Figure 13.29
Figure 13.28
Figure 13.30
14. Mechanisation and Welding Fixtures
2003
14. Mechanisation and Welding Fixtures
188
As the production costs of the metal-working industry are nowadays mainly determined by the costs of labour, many factories are compelled to rationalise their manufacturing methods Designation
movement/ working cycles
examples gas-shielded arc welding TIG GMAW
torch-/ workpiece control
filler wire feeding
workpiece handling
manually
manually
manually
manually
mechanically
manually
mechanically mechanically
manually
fully
manual welding m
v automatic welding
partially
and
mechanised
production
proc-
esses. In the field
partially mechanised welding t fully mechanised welding
by
of
welding
neering
where
consistently mechanically mechanically mechanically
a
engia
good
quality with a maximum productivity is
br-er14-01e.cdr
a must, automation aspects are consequently taken into
Figure 14.1
account.
The levels of mechanisation in welding are stipulated in DIN 1910, part 1. Distinctions are made with regard to the type of torch control and to filler addition and to the type of process sequence, as, e.g., the transport of parts to the welding point. Figure 14.1 explains the four levels of mechanisation.
Figure 14.2. shows manual welding, in this case: manual electrode welding. The control of the electrode and/or the arc is carried out manu-ally. The filler metal (the consumable elecbr-er14-02e.cdr
trode) is also fed manually to the weld-
© ISF 2002
Manual Welding (Manual Electrode Welding)
ing point. Figure 14.2
14. Mechanisation and Welding Fixtures In
189
partially
mechanised welding,
e.g.
gas-
shielded
metal-arc
welding,
the
arc
manipulation is carried out manually, the filler metal addition, however, is executed mechanibr-er14-03e.cdr
cally by means of a wire
feed
Figure 14.3.
Partially Mechanised Welding (Gas-Shielded Metal-Arc Welding)
motor, Figure 14.3
In fully mechanised welding, Figure 14.4, an automatic equipment mechanism carries out the welding advance and thus the torch control. Wire feeding is
realised
by
means of wire feed units.
The
pieces
work-
must
positioned
be
manu-
ally in accordance br-er14-04e.cdr
Fully Mechanised Welding (Gas-Shielded Metal-Arc Welding)
Figure 14.4
with the direction of the
moving
ma-
chine support.
In automatic welding, besides the process sequences described above, the workpieces are mechanically positioned at the welding point and, after welding, automatically trans-ported to the next working station. Figure14. 5 shows an example of automatic welding (assembly line in the car industry).
14. Mechanisation and Welding Fixtures
190 Apart from the actual
welding
de-
vice, that is, the welding source,
power the
filler
metal feeding unit and
the
simple
torch control units, there is a variety of auxiliary
devices
available
which
br-er14-05e.cdr
Automatic Welding (Assembly Line)
facilitate or make Figure 14.5
the welding process at all possible. Figure 14.6 shows
assembly line
a survey of the
welding robot
most
important
machine carrier
assisting devices.
linear travelling mechanism track-mounted welding robots
Before
spindle / sliding head turntable turn-/ tilt table
welding,
the parts are nor-
dollies
mally aligned and
assembly devices
then tack-welded. br-er14-06e.cdr
Figure 14.7 depicts a
simple
welding Figure 14.6
tackjig
for
pipe clamping. The
lower part of the device has the shape of a prism. This allows to clamp pipes with different diameters.
Devices, however, may be significantly more complex. Figure 14.8 shows an example of an assembly equipment used in car body manufacturing. This type of device allows to fix complex parts at several points. Thus a defined position of any weld seam is reproducible.
14. Mechanisation and Welding Fixtures
191 In apparatus engineering
and
tank
construction it is often
necessary
to
rotate the components,
e.g.,
when
welding circumferential
seams.
The
equipment should be as versatile as posbr-er14-07e.cdr
Simple Tack Welding Jig for Welding Circumferential Welds
sible and suit several tank diameters. Figure 14.9
Figure 14.7
shows
three types of turning rolls which fulfil the demands. Figure 1 portal with 2 industrial robots IR 400, equipped with tool change system 2 resting transformer welding tongs 3 depot of welding tongs 4 clamping tool 5 copper back-up bar for car roof welding 6 transformer welding tongs for car roof welding 7 driverless transport system 8 component support frame 9 swivelled support for component support frames 10 resting transformer welding tongs for car boot
br-er14-08e.cdr
top: the rollers are adjustable;
middle: the rollers automatically adapt to the tank diameter; Figure bottom: the roller spacing may be varied by a scissor-like
Figure 17.8
Figure
arrange-
ment.
In general, dollies are motor-driven. This provides also an effortless movement of heavy components, Figure 14.10.
14. Mechanisation and Welding Fixtures
192
set of rollers 2
set of rollers 1 br-er14-09e.cdr
br-er14-10e.cdr
Turning Rolls
Turning Rolls
Figure 14.9
Figure 14.10
A work piece positioner, e.g. a turn-tilt-table, is part of the standard equipment of a robot working station. Figure 14.11 shows a diagrammatic representation of a turntilt-table. Rotations table top rotational axis
gear segment table support tilting axis
around the tilting axis
of
approx.
135° are possible
support
while the turn-table can be turned by 365°. Those types of turn-tables are designed for working
parts
with
br-er14-11e.cdr
weights of just a few kilograms right Figure 14.11
up to several hundred tons.
14. Mechanisation and Welding Fixtures
193
A turn-tilt table with hydraulic adjustment of the tilting and vertical motion as well as chucking grooves for the part fixture is depicted
in
Fig-
ure 14.12.
br-er14-12e.cdr
Turn-Tilt-Table With Hydraulic Adjustment
Figure 14.12
In robot technology the types of turn-tilt-tables - as shown in Figure 14.13 - are gaining importance. Positioners with orbital design have a decisive advantage because the component, when turning around the tilting axis, remains approx. equally distant to the welding robot.
single-column turn-tilt-table table top
table support
orbital turn-tilt-table table top
tilting axis support
tilting axis support rotational axis
rotational axis
© ISF 2002
br-er14-13e.cdr
Turn-Tilt-Tables
Figure 14.13
table support
14. Mechanisation and Welding Fixtures
194
Other types of workpiece positioners are shown in Figure 14.14 – the double column turn-tilt-table and the spindle and sliding holder turn-tilt-table. Those types of positioners are used for special component geometries and allow welding of any seam in the flat and in the horizontal position.
tilting axis
rotational axis table top table support
support
© ISF 2002
br-er14-14e.cdr
Double-Column Turn-Tilt-Table
Figure 14.14 table tops
spindle holder sliding holder
bed way
© ISF 2002
br-er14-15e.cdr
Spindle / Sliding Holder Turntable
Figure 14.15
In the field of welding, special units are designed for special tasks. Figure 14.16 shows a pipe-flange-welding machine. This machine allows the welding of flanges to a pipe. The weld head has to be guided to follow the seam contour.
14. Mechanisation and Welding Fixtures
195
br-er14-16e.cdr
Figure 14.16
Plain plates or rounded tanks are clamped by means of longitudinal jigs for the welding of a longitudinal seam, Figure 14.17. The design and the gripping power are very dependent of the thickness of the plates to be welded.
br-er14-17e.cdr
Figure 14.17
A simple example of a special welding machine is the tractor travelling carriage for submerged-arc welding, Figure 14.18. This device is designed for the application
14. Mechanisation and Welding Fixtures
196 on-site and provides, besides the supply of the filler metal, also the welding speed as well as the feeding and suction of the welding flux.
For the guidance of a welding head and/or welding device, machine supports may be used. Figure 14.19 shows different types of machine supports for welding and cutting. Apart from the translatory and rotary principal axes they are often also equipped with additional axes to allow precise positioning. br-er14-18e.cdr
Tractor for Submerged-Arc Welding
To increase levels of mechanisation of welding processes robots are fre-
Figure 14.18 quently
applied.
Robots are handling
boom main piloting system case
devices which are
pillar
equipped with more
travelling mechanism
than
three
e
d
axes. Figure 14.20
auxiliary piloting system case
kine-
auxiliary piloting system case
matic chains which can be realised by different
cross piloting system case
user-
programmable
describes
c
b
a
br-er14-19e.cdr
combina-
tions of translatory and rotary axes.
Figure 14.19
14. Mechanisation and Welding Fixtures
designation
cartesian robot
cylinder coordinated robot
197
spherical coordinated robot
horizontal knuckle arm robot
vertical knuckle arm robot
arrangement
A
R
x
kinematic schedule
z y
z C
R
D
B
B
C z
C
C
operating space
© ISF 2002
br-er14-20e.cdr
Kinematic Chains
Figure 14.20
The most common design of a trackmounted welding robot is shown in Figure 14.21. The robot depicted here is a hinged-arm robot with six axes. The axes are divided into three principal and three additional axes or hand axes. The wire feed unit and the spool carriers for the wire electrodes are often fixed on the robot. This allows a compact welding design.
br-er14-21e.cdr
Robot Motions
Figure 14.21
14. Mechanisation and Welding Fixtures
198
Varying lever lengths permit the design of robots with different operating ranges. Figure 14.22 shows the operating range of a robot. In the unrestricted operating range the component may be reached with the torch in any position. The restricted operating
range
allows the torch to reach the component
only
certain
positions.
In
the
case
a
sus-
of
pended
arrange-
ment
the
robot
fixing
device
is
shortened thus albr-er14-22e.cdr
lowing a compact design. Figure 14.22
For the completion of a robot welding station workpiece positioners are necessary. Figure 14.23 shows positioner devices where also several axes may be combined. These axes may either turn to certain defined positions or be guided by the robot control and moved synchronically with the internal axes. The complexity and versatility
of
the
axis positions increases
with
number
of
the axes
which participate in the movement. br-er14-23e.cdr
Figure 14.23
14. Mechanisation and Welding Fixtures
199
Movement by means of a linear travelling mechanism increases the operating range of the robot, Figure 14.24. This may be done in ease of stationary as well as suspended arrangement, where there is a possibility to move to fixed end positions or to stay in a synchronised motion with the other movement axes.
br-er14-24e.cdr
Figure 14.24
15. Welding Robots
2003
15. Welding Robots
200
Increased quality requirements for products and the trend to automate production processes along with increased profitability result in the use of industrial robots in modern
manufac-
turing, Figures 15.1 – 15.2. Since robots
have
been
introduced
in
in-
dustry in the 70s, their
most
quently
fre-
fields
of
application ranged from
installation
jobs up to spot welding, and seam Figure 15.1
welding.
The definition says that an industrial robot for gas welding is an universal movement automaton with more than three axes which are user-programmable and may be sensor-controlled. It is equipped with a welding torch and carries out welding jobs.
Core of a modern robot welding cell are one or more seam welding robots of swan neck type. Normally, they have six user-programmable axes; so they can access any point
within
the
working range at any orientation of the welding torch. To
extend
working
their range,
robots may be installed in overhead position. A further extension
of
the
working range can be Figure 15.2
achieved
by
15. Welding Robots
201
installation of the robot onto a linear carriage with Cartesian axes. Such 'external' axes are also user-programmable, Figure 15.3.
To turn the workpiece in the welding-favourable downhand position and to ensure accessibility to any joints, workpiece positioners are used as external axes which are steered by the robot control. Multistation cycle tables are often used to increase profitability of the complete system installation. The operator feeds and removes the welded workpiece on one side, while the robot is welding on the other side.
Figure 15.3
The robot control is the centre of an industrial robot system for arc welding, Figure 15.4. It provides and processes all information for robot mechanics, positioner, welding unit, safety equipment, and external sensors. The robot program transforms information into signals for control of robot-
and
posi-
tioner-mechanics as well
as
power
welding source.
Communication with external systems is possible by a host or master computer.
Figure 15.4
15. Welding Robots
202
Modern industrial robot controls are build as multi-processor controls due to the multitude of parallel calculations and control functions. Figure 15.5 shows the internal structure of such a control. Individual assemblies which are designed for special jobs and equipped with an own micro-processor are linked with the host computer via the system bus. The host controls and coordinates the actions of the components based on the operating system and the robot program. Examples of such assemblies, which are mostly installed on individual printed boards, are e.g. the axes computers. They are responsible for calculation
of
movement and for control of power units of the individual
axes.
To
control the drive motors, two interconnected control loops per axis are available
which
control speed and position of each
Figure 15.5
axis.
Further
assemblies
control the display screen, the manual programming
unit
(PHG);
these
as-
are
re-
semblies sponsible
for
communication with the welding power source,
external
sensors, and peripheral units via digital Figure 15.6
15. Welding Robots
203
and analogue in- and outputs and field bus systems. Or they complete the data transmission with external control systems. To reduce downtimes in the case of malfunction, some robot controls can be connected via internet with telediagnosis systems of the robot manufacturer to support service personnel during troubleshooting and commissioning.
Programming of welding robots can be carried out in different ways which are distinguished in On-Line (programming at the robot) and Off-Line (programming out of the robot cell), Figure 15.6.
The robot is manually guided along the later track with decoupled drives during PlayBack programming. The path of the track is recorded and transformed into a corresponding robot control program. This procedure is preferably used for painting jobs.
A common technique to program a robot is the Teach-In procedure. During Teach-In programming, with the help of the manual programming unit, the welding torch is moved to notable points of the groove to be welded which are stored with information about position and orientation. In addition, track parameters must be entered, like e.g. type of movement and speed or welding parameter sets.
During sensor supported Teach-In programming, the path progress through some typical points is only roughly indicated. Then the accurate path is picked-up by sensors
and
auto-
matically
calcu-
lated in the robot steering
control.
Afterwards
the
movement
pro-
gram is supplemented
by
additional information
about
e.g.
welding parameter sets. Figure 15.7
15. Welding Robots
204
Textual programming belongs to mixed procedures. The sequence program in form of a text file is created on an external computer and is then transmitted to the robot steering control, Figure 15.7. The recording of the position of points is carried out in the same way as with Teach-In programming: moving into position and recording.
Macro-programming is also regarded as a mixed method which shortens programming time at the robot, Figure 15.8. Macros are structured processing sequences which are created online to fulfil working functions and which can be repeated for further similar working functions. Geometry macros contain information about torch guidance to produce certain joints or joint sections. Welding
technol-
ogy parameters for individual
welding
situations
are
summarised welding This
Figure 15.8
in
macros.
applies
for
torch
positioning,
torch
inclination,
relative position of beads to root and welding
parame-
ters.
Using a collection (can
be
created
online or offline) of such macros, the programming time can be shortened for workpieces with often Figure 15.9
repeated
15. Welding Robots
205
welding jobs, e.g. steel construction when welding stiffeners and head plates Using offline programming practice, the programming work is shifted out from the producing robot cell. This avoids unproductive stoppages and allows for economicviable, limited number of pieces to be reduced. During textual programming, the 3-dimensional point coordinates and torch orientations are entered into an external computer in a manufacturer-specific program language. To achieve a complete program sequence, each instruction must be entered individually.
The graphical offline programming uses CAD data for modelling the complete robot working cell and parts to be welded. Planning of the path is carried out with CAD functions directly at the workpiece which is displayed on a screen. In most cases, the programming systems provide a graphical simulation of the movement, e.g. to check for collisions between
torch
workpiece,
and
Figure
15.9. For the following transformation of the program into the robot control, a calibration between
model
and physical robot working cell is required. Figure 15.10
In the case of knowledge-based offline programming, the operator is supported by integrated expert systems when it comes to creation of robot welding programs, e.g. for determination of job-specific welding parameters. However, checking and adapting the program must be carried out by the operator. Modern robot controls provide the programmer with some functions for movement control and for modification of program sequence, Figure 15.10. PTP movement (point to point) serves to move the robot in the space. All axes are controlled in such
15. Welding Robots
206
a way that they reach their set-point at the same time. Thereby the actual path of the torch depends on kinematics of the robot and on current position of the axes.
A linear interpolation (CP procedure, continuous Path), Figure 15.11, is used for accurate movement along a straight line, e.g. movement to weld start point or welding. The active point of the tool 'arc' (ToolCentre-Point, TCP) is moved along a straight tween
line
be-
two
grammed
propoints,
adapting torch angle and torch inclination between the two points. Figure 15.11
Circles and graduated circles are entered by means of circle interpolation programs, Figure 15.12. Then the orientation of the torch can be adapted through turning the knuckle axis or 6th axis of the robot and the value of spill-weld at the end of the seam can be indicated.
Speed of the torch is
user-
programmable and, if required, can be superimposed
by
an
To
oscillation.
control the program run, commands are available
for:
re-
peated loops, conFigure 15.12
15. Welding Robots
207
ditional and unconditional program jumps, waiting periods, waiting for inputs, and working with sub-programs. The software of modern seam welding robots contains – as special functions – 3dimansional transfor-mations and mirroring of programs and partial programs, palletising
functions,
processing sensor data
and
com-
mands for communication with other robot
controls
(Master/Slave
op-
eration) as well as with external computers, 15.13. Figure 15.13
Figure
16. Sensors
2003
16. Sensors
208
The welding process is exposed to disturbances like misalignment of workpiece, inaccurate preparation, machine and device tolerances, and proess disturbances, Figure 16.1.
The manual welder notices them by eyesight and corrects them manually according to strategies learned and gained by experience. To record process irregularities and path deviations, a fully mechanised welding plant requires sensors providing control signals which are then used in accordance with implemented rules. Using corresponding control elements, the control loop is closed for the welding process.
Scopes of duty of the sensors is finding the weld start point
and
seam
tracking. In addition, with the help of information joint
about geometry,
process parameters can
be
adapted
online and offline. The ideal sensor for
Figure 16.1
a robot application should measure the welding
point
(avoidance of tracking
misalignment),
detect in advance (finding
the
start
point of the seam, recognising ners, collisions)
coravoiding and
should be as small Figure 16.2
16. Sensors
209
as possible (no restriction in accessibility). The ideal sensor which combines all three requirements, does not yet exist, therefore one must select a sensor which is suitable for
the
welding
individual job.
Fig-
ure 16.2
shows
different
sensor
principles used in welding ing.
engineer-
The
most
frequently used systems in practice are tactile, optical, and arc based sensor systems with mechanical
arc
Figure 16.3
adjustment.
With tactile scanning systems, the simplest type of scanning is a mechanical sensor. Pins, rollers, balls, or similar devices may be used as sensors.
Such scanning systems show a long distance between sensor and torch, the application range is limited. Only grooves with large dimensions and relatively straight seam path can be scanned with these systems. Figure 16.3 shows some examples of different groove geometries.
Tactile sensors can recognise 3dimensional offsets of the workpiece. Figure 16.4
16. Sensors
210
Through scanning of three levels the 3-dimensional point of intersection can be calculated and the robot program for correcting the deviation can be shifted accordingly thus finding the start point of the weld. In this case, the gas nozzle of the torch serves as a sensor, Figure 16.4, which is charged with electrical tension. As soon as the torch touches the workpiece, a current flows, which is then taken by the robot control as a signal for obtaining the level to be scanned.
Inductive sensors are graded as non-contact measurement systems. Due to their function principle, they can be applied for metallic and electrically conductive materials. The simplest type is a ring coil. If alternating current flows though the coil, ,a magnetic field is generated close to the workpiece. When the coil approaches the workpiece surface, the magnetic field weakens. Figure 16.5 shows the distancedependent electrical
signal.
Such
simple sensors are used to recognise the workpiece position.
Using
several
distance
sensors,
also
welding
a
groove
can be scanned. Figure 16.5 With multi-coil arrangements in one sensor, the position of the welding groove, the angle between sensor and workpiece surface and the distance can be recorded. Figure 16.6 shows a principle arrangement. A transmitter coil generates an magnetically alternating field which induces
alternating currents in the two receiver coils. In the undisturbed case, these currents are phase-shifted by 180° and neutralise each other. If the sensor is moved crosswise to the groove, magnetical asymmetries will occur in the scanning area, which
16. Sensors
211
will show in the presented signal shape. The output signal will be zero, if the coils are positioned exactly above the centre of the groove. The radar sensor in Figure 16.6 uses Doppler's effect to generate a signal. Here the phase difference between transmitter signal and receiving signal is evaluated. A mathematical process transforms such signals into distance values. To record the position and the depth of the groove, the sensor must be continuously moved along the seam. Radar sensors form a so called radar baton, which is focussed onto a measurement spot of about 0,7 mm diameter for this application. Figure 16.6 shows the sensor signal, which
represents
the relative movement
along
the
workpiece. At the moment, the characteristic values of the
weld
groove
can be determined with a resolution in the range of 1/10 mm. Figure 16.6 Arc sensors evaluate the continuous change welding
of
the
current
with a change of the contact tip-towork distance, Figure 16.7. A signal for side control of the torch is determined
by
measurement and Figure 16.7
16. Sensors
212
subtraction of the currents on the flanks of a groove. A comparison between actual welding current and programmed rated current provides a signal for distance control of the welding torch. To let this sensor method work, a divergence of the arc or the use of a second arc is required.
To realise this principle, there are numerous possibili-ties. Figure 16.8 shows some variants of signal recording. The most frequently used method is a mechanical oscillation of the welding torch, which is carried out by a rotor movement with an oscillation frequency up to 5 Hz. The second method is mainly used with submerged
arc
welding. Both wires are aligned crossways
to
direction
welding and
the
difference of the two currents
is
evalu-
ated. Figure 16.8
Magnetic fields can diverge only the arc itself. The advantage of this method is a high divergence frequency of about 15 Hz. A disadvantage is the size of the electromagnets and the limited accessibility to the workpiece. The last variant of an arc sensor incorporates a mechanical rotation of the welding wire. In this case, the divergence frequency of the arc can reach up to 30 Hz.
The signal recording is continuous during the movement. In this way, information about orientation of the torch and groove width is also provided. The arc sensor principle is limited to groove shapes with clear flanks. Together with the tactile torch gas nozzle sensor, it provides a frequently used combination for seam finding and seam tracking during robot welding.
16. Sensors
213
Optical sensors can be used for a great number of jobs. The easiest method is the recognition
of
the
radiation
intensity,
which
reflected
is
during welding. E.g. with laser beam welding, this is carried
out
recording flected
through the
relaser
radiation with simple sensors for control of
penetration
depth, Figure 16.9.
Figure 16.9
The procedure is based on the line-up between the degree of reflection and shaft relation (penetration depth/focus position) of the capillary. The amount of backreflection of the laser beam power is measured, which due to multi-reflection is not absorbed by the workpiece. Changes of penetration depth due to modified laser power or a shifted focus position can be identified by the signal of reflected laser power and can be used for control of the penetration depth. However, optical sensors can also be used for measuring geometrical values. Such information may be used for finding the start point of a seam, for seam tracking, and for identification of groove profile. The two last mentioned functions provide the possibility to use the information for filling rate control and/or quality control.
Geometry-measuring optical sensors are normally external systems, which are positioned in front of the torch as a leading element. It is practical to equip the sensor with additional axes, because both, torch and sensor, must be moved along the groove. Without additional axes, a robot would be limited in its accessibility to the workpiece and in its working range. Another problem is the tremendous effort to introduce the control-technical integration into the robot control. Among other things, information must be exchanged in real time.
16. Sensors
214
Most of geometry-measuring sensors use the triangulation principle or a variant of this measurement procedure. The triangulation measurement procedure provides information about the distance to the workpiece surface. A light spot is projected onto the workpiece surface and displayed to a line-type receiver element under a certain angle. With distance changes emerge corresponding positions on the receiver element, Figure 16.10. Sensors which use this triangulation principle are applied for recognition of workpiece position and for offline seam finding. Figure 16.10 Both, the laser scanner and the light-section procedure are based on the triangulation measurement principle. With the laser scanner, Figure 16.11, this principle is complemen-ted by an oscillating axis in parallel to the groove axis. The measurement of a sequence of distances along a line becomes possible and provides a 2-dimensional
re-
cord and evaluation of the groove contours.
Sensors as part of the
light-section
procedure,
also Figure 16.11
16. Sensors
215
provide information about the 2-dimensional position of the groove. As a function of this system, one or more light lines are projected onto the workpiece surface and displayed to a CCD matrix under a certain angle, Figure 16.12. In contrast to scanning, information about the groove profile is provided by taking a picture scene. Using sensors, it is pssible to obtain additional 3-dimensional information through evaluation of more, in succession taken, while the camera moves over the grooves. Systems, which generate their information through a projection of several light lines, provide additional information about the path of the seam and the orientation of the sensor related to the workpiece surface. Both,
scanning
systems and sensors based on the light section procedure, can be used for recognition of the welded seam to make
an
automised
quality
control of the outer weld
characteris-
tics possible. Figure 16.12
Another
optical
measurement
prin-
ciple uses, similar to human
sight,
the
stereo procedure to record
geometry
information the
weld
Two optics the
across groove.
independent photograph interesting Figure 16.13
16. Sensors
216
groove area and displays them onto two image converter elements (CCD-lines or CCD-matrix). Based on the corresponding image points in both picture scenes, the 3dimensional position of object points is evaluated. Figure 16.13 shows the measurement principle, which uses CCD lines as image converter elements, and idealised signals for generating information. The grey scale drop in the signal is ideally used as corresponding image area, which occurs with butt welds due to different reflection intensity between workpiece surface and gap. Both, the lateral position of the groove and the distance to the sensor can be determined by evaluating the centre positions of both signal drops. The width of the groove is taken from the width of the signal drop.
Optical sensors may also be used for geometrical recognition of the weld pool, to adapt process parame-ters in the case of possible deviations. Figure 16.14 depicts such a system for use with laser beam welding. The welding process is monitored by a CCD camera through a filter system. An optical filter allows to observe the weld pool surface without disturbing effects of the plasma in the near infrared spectrum. Picture data are transferred to an image processing computer which measures the geometry of the weld pool. Geometry data contain information which is used online for control of the welding
process.
Among
others,
penetration
depth
and focus position can be controlled. The
system
also
provides the recognition of protrusionwelded joints and welding defects like e.g.
molten
ejections.
pool Figure 16.14
16. Sensors
217
During electron beam welding, the beam is in combination with a detector used for both, to carry out a seam tracking and a monitoring of the welded seam. For this, the beam can be diverged as well as bent, Figure 16.15. Backscattered electrons are recognised by a special detector and converted into grey values. The line or area surface scanning by the spotted electron beam provides a progressive series of greys across the scanned line or area. During electron beam welding, these signals can be used for seam tracking by scanning an edge which is parallel to the
groove.
The
area-type scanning provides the possibility
for
observing
the
welded
seam
or
the focus position. Figure 16.15
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225 Neuere Entwicklungen beim Plasmaschneiden Trennen und Fügen, 1985, Heft 15, S. 55-58 Ruckdeschel, W. Plasmaheißdraht-Auftragschweißen – Ein neues Plattierungsverfahren DVS-Bericht Band 23/1972 Ruge, J. Handbuch der Schweißtechnik, Bd. II, Verfahren und Fertigung Springer-Verlag, Berlin Heidelberg New York 1980 Schäfer, P. Industrielle Anwendungen von Festkörperlasern Laser und Optoelektronik, 2/1988 Schellhase, M. Der Schweißlichtbogen – ein technologisches Werkzeug VEB Verlag Technik, Berlin 1985 Schiller, S. et al. Elektronenstrahltechnologie Wissenschaftliche Verlagsgesellschaft mbH, Stuttgart, 1977 Schmidt, H. u. K. Ludewig Hochleistungs-Festkörperlaser Laser und Optoelektronik, 2/1988 Schultz, H. Elektronenstrahlschweißen DVS-Verlag, Düsseldorf, 1989 Seiler, P. Schweißen mit YAG-Laser Feinwerktechnik & Messtechnik, 96 (1988) 7-8 SOUDOMETAL Firmenprospekt Taylor D.S. u. C.E. Thornton High Deposition Rate Submerged-Arc Welding Welding Review, Aug. 1989 Tong S. u. Z. Ding Effect Of Plasma Spraywelding Technology On Dilution Wuhan (China) 1985 Tradowsky, Klaus Laser: Grundlagen, Technik, Basisanwendungen, Kamprath-Reihe Technik
Literature
226
Literature
Vogel-Uerlag Würzburg 1988 Wahl, W. Auftragschweißen – Standzeitverlängerung durch gezielten Werkstoffeinsatz und optimale Schweißverfahren Schweißen und Schneiden 6/79 Yamamoto, H. Recent Trends in Low Current Airplasma Cutting Welding International 55, 1987, S. 35-43
ISF – Welding and Joining Institute RWTH – Aachen University
Lecture Notes
Welding Technology 2 Welding Metallurgy
Prof. Dr. –Ing. U. Dilthey 2005
Table of Contents
Chapter
Subject
Page
1.
Weldability of Metals
3
2.
TTT - Diagrams
8
3.
Residual Stresses
4.
Heat Treatment and its Function During Welding
5.
21 31
Welding Plain and Low Alloy Steels
44
6.
Welding High Alloy Steels
70
7.
Welding of Cast Materials
89
8.
Welding of Aluminium
96
9.
Welding Defects
108
Testing of Welded Joints
126
10.
1. Weldability of Metals
1. Weldability of Metals
4
DIN 8580 and DIN 8595 classify welding into production technique main group 4 "Joining“, group 3.6 "Joining by welding“, Figure 1.1.
Figure 1.1
Weldability of a component is determined by three outer features according to DIN 8528, Part 1. This also indicates whether a given joining job can be done by welding, Figure 1.2.
Figure 1.2
1. Weldability of Metals
5 Material influence on weldability, i.e. welding suitability, can be detailed for a better understanding in three subdefinitions, Figure 1.3.
The chemical composition of a material and also its metallurgical properties are mainly set during its production, Figure 1.4. They have a very strong influence on the physical characteristics of the material. Process steps on steel manufacturing, shown in Figure 1.4, are the essential steps on the way to a processible and usable material. During manufacture, the requested chemical composition (e.g. by alloying) and metallurgical properties (e.g. type of teeming) of the steel are obtained. Figure 1.4
Another modification of the material beha viour takes place during subsequent treatment, where the raw material is rolled to processible semi-finished goods, e.g. like strips, plates, bars, profiles, etc.. With the rolling process, materialtypical transformation processes, hardening and precipitation processes are used to adjust an optimised
material
(see chapter 2).
characteristics
Figure 1.3
1. Weldability of Metals
6
A survey from quality point of view about the influence of the most important alloy elements to some mechanical and metallurgical properties is shown in Figure 1.5.
Figure 1.5
Figure 1.6 depicts the decisive importance of the carbon content to suitability of fusion welding of mild steels. A guide number of flawless fusion weldability is a carbon content of C < 0,22 %. with higher C contents, there is a danger of hardening, and welding becomes only possible by observing special precautions (e.g. pre- and post-weld heat treatment).
Figure 1.6
1. Weldability of Metals
7
In addition to material beha viour, weldability is also essentially determined through the design of a component. The influence of the design is designated as welding safety, Figure 1.7.
Figure 1.7
The influence of the manufacturing process to weldability is called welding possibility, Figure 1.8. For example, a pre-
and
post-weld
heat
treatment is not always possible, or grinding the weld surface
before
welding
the
subsequent pass cannot be carried out (na rrow gap welding).
Figure 1.8
2. TTT - Diagrams
2. TTT – Diagrams
9
An essential feature of low alloyed ferrous materials is the crystallographic transformation of the body-centred cubic lattice which is stable at room α -Iron body-centered
temperature (α-iron, ferritic structure)
to
the
γ -Iron face-centered
face-
centred cubic lattice (γ-iron, austenitic structure), Figure 2.1. The temperature, where this transformation occurs, is not constant but depends on Lattice constant 0.364 nm at 900 °C
Lattice constant 0.286 nm at room temperature
factors like alloy content,
br-eI-02-01.cdr
crystalline
structure,
ten-
sional status, heating and cooling rate, dwell times,
Figure 2.1
etc.. In order to be able to understand the basic processes it is necessary to have a look at the basic processes occuring in an idealized binary system. Figure 2.2 shows the state of a binary system with complete solubility in the liquid and solid state. If the melting of the L1 alloy is cooling down, the first crystals of the composition c1 are formed with reaching the temperature T1. These crystals are depicted as mixed crystal α, since they consist of a compound of the components A (80%) and of B (20%). Further, a melting with the composition c0 is present at the temperature T1. L1
L1
With dropping temperature,
S TsA
riched with component B,
T2
T1
Li (liquidus line, up to point
3 2
4 5
Temperature T
following the course of line
1
Li So
TsB
Temperature T
the remaining melt is en-
α - ss
4). In parallel, always new
a
b
and B richer α-mixed crystals are forming along the
A (Ni) br-eI-02-02.cdr
connection line So (solidus line, points 1, 2, 5). The dis-
Figure 2.2
c1
c2
c0
c3
Concentration c
c4
B (Cu)
Time t
2. TTT – Diagrams
10
tribution of the components A and B in the solidified structure is homogeneous since concentration differences of the precipitated mixed crystals are balanced by diffusion processes. The other basic case of complete solubility of two components in the liquid state and of complete insolubility in the solid state shows Figure 2.3 If two components are completely insoluble in the solid state, no mixed crystal will be formed of A and B. The two liquidus lines Li cut in point e which is also designated as the eutectic point. The isotherm Te is the eutectic line. If an alloy of free composition solidifies according to Figure 2.3, the eutectic line must be cut. This is the temperature (Te) of the eutectic transformation: S → A+B (T = Te = const.). This means that the melt at a constant temperature Te dissociates in A and B. If an alloy of the composition L2 solidifies, a purely eutectic structure results. On account of the eutectic reaction, the temperature of the alloy remains constant up to the completed transformation (critical point) (Figure 2.2). Eutectic L1
L2
L1
TsA
1 TsB
So Te
Li
Li
S+A
S+B
a
orientation
Temperature T
2
are
normally fine-grained and show
S 2’
L2
structures
characteristic between
the
constituents. The alloy L1 will consist of a compound
3 4
of alloy A and eutectic alloy A+E
E
B+E
E in the solid state. A
c1
ce
Concentration c
B
Time t
You can find further infor-
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mation on transformation behaviour in relevant specialist literature.
Figure 2.3 The definite use of the principles occurs in the iron-iron carbide diagram. Transformation behaviour of carbon containing iron in the equilibrium condition is described by the stable phase diagram iron-graphite (Fe-C). In addition to the stable system Fe-C which is specific for an equilibrium-close cooling, there is a metastable phase diagram iron cementite (Fe-Fe3C). During a slow cooling, carbon precipitates as graphite in accord with the stable system Fe-C,
2. TTT – Diagrams
11
while during accelerated cooling, what corresponds to technical conditions, carbon precipitates as cementite in agreement with the metastable system (Fe-Fe3C). Per definition, iron carbide is designated as a structure constituent with cementite although its stoichiometric composition is identical (Fe3C). By definition, cementite and graphite can be present in steel together or the cementite melt + δ - solid solution
can decompose to iron δ− solid sol.
and graphite during heat
δ −+γ−
Fe3C (cementite)
solid sol.
melt + austenite
treatment of carbon rich
austenite + graphite austenite + cementite
formation of cementite is
ledeburite
damentally valid that the
melt + cementite
austenite
Temperature °C
alloys. However, it is fun-
austenite + ferrite
encouraged with increas-
ferrite
ferrite + graphite ferrite + cementite
perlite
ing cooling rate and decreasing carbon content.
stable equilibrium metastable equilibrium
Mass % of Carbon
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In a double diagram, the
Stable and Metastable Iron-Carbon-Diagram
stable system is shown by a dashed, the metastable
melt + graphite
melt
Figure 2.4
by a solid line, Figure 2.4. The metastable phase diagram is limited by the formation of cementite with a carbon content of 6,67 mass%. The strict stoichiometry of the formed carbide phase can be read off at the top X-coordinate of the molar carbon content. In accordance with the carbon content of Fe3C, cementite is formed at a molar content of 25%. The solid solutions in the phase fields are designated by Greek characters. According to convention, the transition points of pure iron are marked with the character A - arrêt (stop point) and distinguished by subjacent indexes. If the transition points are determined by cooling curves, the character r = refroidissement is additionally used. Heat-up curves get the supplement c - chauffage. Important transition points of the commercially more important metastable phase diagram are:
-
1536 °C: solidification temperature (melting point) δ-iron,
-
1392 °C: A4- point γ- iron,
-
911 °C: A3- point non-magnetic α- iron,
with carbon containing iron: -
723 °C: A1- point (perlite point).
2. TTT – Diagrams
12
The corners of the phase fields are designated by continuous roman capital letters. As mentioned before, the system iron-iron carbide is a more important phase diagram for technical use and also for welding techniques. The binary system iron-graphite can be stabilized by an addition of silicon so that a precipitation of graphite also occurs with increased solidification velocity. Especially iron cast materials solidify due to their increased silicon contents according to the stable system. In the following, the most important terms and transformations should be explained more closely as a case of the metastable system. The transformation mechanisms explained in the previous sections can be found in the binary system iron-iron carbide almost without exception. There is an eutectic transformation in point C, a peritectic one in point I, and an eutectoidic transformation in point S. With a temperature of 1147°C and a carbon concentration of 4.3 mass%, the eutectic phase called Ledeburite precipitates from cementite with 6,67% C and saturated γ-solid solutions with 2,06% C. Alloys with less than 4,3 mass% C coming from primary austenite and Ledeburite are called hypoeutectic, with more than 4,3 mass% C coming from primary austenite and Ledeburite are called hypereutectic.
If an alloy solidifies with less than 0,51 mass percent of carbon, a δ-solid solution is formed below the solidus line A-B (δ-ferrite). In accordance with the peritectic transformation at 1493°C, melt (0,51% C) and δ-ferrite (0,10% C) decompose to a γ-solid solution (austenite).
The transformation of the γ-solid solution takes place at lower temperatures. From γ-iron with C-contents below 0.8% (hypoeutectoidic alloys), a low-carbon α-iron (pre-eutectoidic ferrite) and a fine-lamellar solid solution (perlite) precipitate with falling temperature, which consists of α-solid solution and cementite. With carbon contents above 0,8% (hypereutectoidic alloys) secondary cementite and perlite are formed out of austenite. Below 723°C, tertiary cementite precipitates out of the α-iron because of falling carbon solubility.
The most important distinguished feature of the three described phases is their lattice structure. α- and δ-phases are cubic body-centered (CBC lattice) and γ-phase is cubic facecentered (CFC lattice), Figure 2.1.
2. TTT – Diagrams
13
Different carbon solubility of solid solutions also results from lattice structures. The three above mentioned phases dissolve carbon interstitially, i.e. carbon is embedded between the iron atoms. Therefore, this types of solid solutions are also named interstitial solid solution. Although the cubic face-centred lattice of austenite has a higher packing density than the cubic body-centred lattice, the void is bigger to disperse the carbon atom. Hence, an about 100 times higher carbon solubility of austenite (max. 2,06% C) in comparison with the ferritic phase (max. 0,02% C for α-iron) is the result. However, diffusion speed in γ-iron is always at least 100 times slower than in α-iron because of the tighter packing of the γ-lattice.
Although α- and δ-iron show the same lattice structure and properties, there is also a difference between these phases. While γ-iron develops of a direct decomposition of the melt (S → δ), α-iron forms in the solid phase through an eutectoidic transformation of austenite (γ → α + Fe3C). For the transformation of non- and low-alloyed steels, is the transformation of δferrite of lower importance, although this δ-phase has a special importance for weldability of high alloyed steels. Unalloyed steels used in industry are multi-component systems of iron and carbon with alloying elements as manganese, chromium, nickel and silicon. Principally the equilibrium diagram Fe-C applies also to such multi-component systems. Figure 2.5 shows a schematic cut through the three phase system Fe-M-C. During precipitation, mixed carbides of the general composition M3C develop. In contrast to the binary system Fe-C, is the three phase system Fe-M-C characterised by a
Ac3
temperature interval in the Ac1e
three-phase field α + γ + M3C. The beginning of the transformation of α + M3C to γ is marked by Aclb, the end by Acle. The indices b and e mean the beginning
br-eI-02-05.cdr
Description of the Terms Ac1b, Ac1e, Ac3
and the end of transformation.
Figure 2.5
2. TTT – Diagrams
14
The described equilibrium diagrams apply only to low heating and cooling rates. However, higher heating and cooling rates are present during welding, consequently other structure types develop in the heat affected zone (HAZ) and in the weld metal. The struc°C
ture transformations during heating and cooling are described
by
transformation
diagrams, where a temperature change is not carried out close to the equilibrium, s
but
at
different
heating br-eI-02-06.cdr
and/or cooling rates. A
representation
transformation
of
TTA Diagram for Isothermal Austenitization
the
processes
Figure 2.6 during isothermal austenitizing shows Figure
ASTM4; L=80µm
2.6. This figure must be read exclusively along
ASTM11; L=7µm
the time axis! It can be recognised that several transformations during isothermal austenitizing occur with e.g. 800°C. Inhomogeneous austen20µm
20µm
ite means both, low carbon containing austenite is formed in areas, where ferrite was present before transformation, and carbon-rich austenite is formed in areas during transforma-
Temperature
tion, where carbon was present before transformation. During sufficiently long annealing times, the concentration differences are balanced by diffusion, the border to a homogeneous austenite is passed. A growing of the Time br-er02-07.cdr
© ISF 2002
austenite grain size (to ASTM and/or in µm) can here simultaneously be observed with longer annealing times.
Figure 2.7
2. TTT – Diagrams
15
The influence of heating rate on austenitizing is shown in Figure 2.7. This diagram must only be read along the sloping lines of the same heating rate. For better readability, a time pattern was added to the pattern of the heating curves. To elucidate the grain coarsening during austenitizing, two microstructure photographs are shown, both with different grain size classes to ASTM. Figure 2.8 shows the relation between the TTA and the Fe-C diagram. It's obvious that the Fe-C diagram is only valid for infinite long dwell times and that the TTA diagram applies only for one individual alloy. Figure 2.9 shows the dif-
Ac3
ferent
time-temperature
passes during austenitizing
Ac1e
and
Ac1b
subsequent
cooling
down. The heating period is composed of a continuous and an isothermal section. br-eI-02-08.cdr
Dependence Between TTA-Diagram and the Fe-M-C System
During cooling down, two different ways of heat con-
Figure 2.8
trol can be distinguished: 1. : During continuous temperature
Ac3
continuous Ac1e
control
a
cooling is carried out with a constant cooling rate out of
Ac1b
the area of the homogeneisothermal
ous and stable austenite down to room temperature. 2.
:
During
temperature
isothermal control
a
quenching out of the area
br-eI-02-09.cdr
Heating and Cooling Behaviour With Several Heat Treatments
of the austenite is carried out into the area of the me-
Figure 2.9
2. TTT – Diagrams
16
the area of the homogeneous and stable austenite down to room temperature. 2. : During isothermal temperature control a quenching out of the area of the austenite is carried out into the area of the metastable austenite (and/or into the area of martensite), followed by an isothermal holding until all transformation processes are completed. After transformation will be cooled down to room temperature.
Figure
2.10
shows
the
time-temperature diagram of a isothermal transformation of the mild steel Ck 45. Read such diagrams only along the time-axis! Below the Ac1b line in this figure, there is the area of the metastable austenite, marked with
an
A.
The
areas
marked with F, P, B, und M represent areas where ferFigure 2.10
rite, perlite, Bainite and martensite are formed. The
lines which limit the area to the left mark the beginning of the formation of the respective structure. The lines which limit the area to the right mark the completion of the formation of the respective structure. Because the ferrite formation is followed by the perlite formation, the completion of the ferrite formation is not determined, but the start of the perlite formation. Transformations to ferrite and perlite, which are diffusion controlled, take place with elevated temperatures, as diffusion is easier. Such structures have a lower hardness and strength, but an increased toughness.
Diffusion is impeded under lower temperature, resulting in formation of bainitic and martensitic structures with hardness and strength values which are much higher than those of ferrite and perlite. The proportion of the formed martensite does not depend on time. During quenching to holding temperature, the corresponding share of martensite is spontanically formed. The present rest austenite transforms to Bainite with sufficient holding time. The right
2. TTT – Diagrams
17
detail of the figure shows the present structure components after completed transformation and the resulting hardness at room temperature. Figure 2.11 depicts the graphic representation of the TTT diagram, which is more important for welding techniques. This is the TTT diagram for continuous cooling of the steel Ck 15. The diagram must be read along the drawn cooling passes. The lines, which are limiting the individual areas, also depict the beginning and the end of the respective transformation. Close to the cooling curves, the amount of the formed structure is indicated in per cent, at the end of each curve, there is the hardness value of the structure at room temperature.
Figure 2.12 shows the TTT diagram of an alloyed steel containing
approximately
the same content of carbon as the steel Ck 15. Here you can see that all transformation
processes
are
strongly postponed in relation to the mild steel. A
Figure 2.11
completely
marte nsitic
transformation
is
carried
out up to a cooling time of about 1.5 seconds, compared with 0.4 seconds of Ck 15. In addition, the completely diffusion controlled transformation processes of the perlite area are postponed to clearly longer times.
The hypereutectoid steel C 100
behaves completely
different, Figure 2.13. With Figure 2.12
this carbon content, a pre-
2. TTT – Diagrams
18 eutectoid ferrite formation cannot still be carried out (see also Figure 2.3). The term of the figures 2.9 to 2.11 "austeniti zing temperature“ means the temperature, where the workpiece transforms to an auste nitic microstructure in the course of a heat treatment. Don’t mix up this temperature with the AC3 temperature, where above it there is only pure auste nite. In addition you can see that only martensite is formed from the austenite, provided that the cooling rate is sufficiently
high,
a
formation
of
any
other
microstructure is completely depressed. With this type of transformation, the steel gains the highest hardness and strength, but loses its toughness, it embrittles. The slowest cooling rate where such a transformation happens, is Figure 2.13
Figure 2.14
called critical cooling rate.
Figure 2.15
2. TTT – Diagrams
19
Figure 2.14 shows schematically how the TTT diagram is modified by the chemical composition of the steel. The influence of an increased austenitizing temperature on transformation beha viour shows Figure 2.15. Due to the higher hardening temperature, the grain size of the austenite is higher (see Figure 2.6 and 2.7).
This grain growth leads to an extension of the diffusion lengths which must be passed during the transformation. As a result, the "noses" in the TTT diagram are shifted to longer times. The lower part of the figure shows the proportion of formed
martensite
and
Bainite depending on cooling time. You can see that Figure 2.16
with
higher
austenitizing
temperature the start of Bainite formation together with the drop of the martensite proportion is clearly shifted to longer times. As Bainite formation is not so much impeded by the coarse austenite grain as with the completely diffusion controlled processes of ferrite and perlite formation, the maximum Bainite Figure 2.17
proportion
is
increased
from about 45 to 75%.
2. TTT – Diagrams
20
Due to the strong influence of the austenitizing temperature to the transformation behaviour of steel, the welding technique uses special diagrams, the so called Welding -TTT-diagrams.
They are recorded following the welding temperature cycle with both, higher austenitizing temperatures (basically between 950° and 1350°C) and shorter a usteniti zing times. You find two examples in Figures 2.16 and 2.17.
Figure 2.18 proves that the iron-carbon diagram was developed as an equilibrium diagram for infinite long cooling time and that a TTT diagram applies always only for one alloy.
Figure 2.18
3. Residual Stresses
3. Residual Stresses
22
The emergence of residual stresses can be of very different nature, see three
pressure
tension
examples in Figure 3.1. Figure
grinding disk
3.2
details
the
causes of origin. In a protension
pressure
duced workpiece, material-
weld
, production-, and wearcaused residual stresses are overlaying in such a
© ISF 2002
br-eI-03-01e.cdr
way that a certain condition
Various Reasons of Residual Stress Development
of residual stresses is cre-
Figure 3.1
ated. Such a workpiece shows in service more or
less residual stresses, and it will never be stress-free! Figure 3.3 defines residual stresses of 1., 2., and 3. type. This grading is independent from the origin of the residual stresses. It is rather based on the three-dimensional extension of the stress conditions. Based on this definition, FigAnalysis of Residual Stress Development
ure 3.4 shows a typical distrirelevant material
bution of residual stresses. Residual
stresses,
which
build-up around dislocations
wear
production
e.g. polyphase systems, non-metallic inclusions, grid defects
and other lattice imperfections
mechanical
thermal
chemical
e.g. partial-plastic deformation of notched bars or close to inclusions, fatigue strain
e.g. thermal residual stresses due to operational temperatur fields
e.g. H-diffusion under electro-chemical corrosion
(σIII), superimpose within a grain causing stresses of the 2
nd
type and if spreading
forming
deforming
separating
joining
plating
e.g. thermal residual stresses
residual stresses due to inhomogenuous deformationanisotropy
residual stresses due to machining
residual stresses due to welding
layer residual stresses
changing material characteristics induction hardening, case hardening, nitriding
around several grains, bring © ISF 2002
br-eI-03-02e.cdr
out residual stresses of the 1
st
Development of Residual Stresses
type. The
formation
of
residual
stresses in a transition-free
Figure 3.2
3. Residual Stresses
23
steel cylinder is shown in Figures 3.5. and 3.6. During water quenching of the homogeneous heated cylinder, the edge of the cylinder cools down faster than the core. Not before 100 seconds have elapsed is the temperature across the cylinder's cross section again
s III
tension s
General Definition of the Term ‘Residual Stresses’
Residual stresses of the I. type are almost homogenuous across larger material areas (several grains). Internal forces related to residual stresses of I. type are in an equilibrium with view to any cross-sectional plane throughout the complete body. In addition, the internal torques related to the residual stresses with reference to each axis disappear. When interfering with force and torque equilibrium of bodies under residual stresses of the I. type, macroscopic dimension changes always develop.
s II sI
+
0
x
-
y
Residual stresses of the II. type are almost homogenuous across small material areas (one grain or grain area). Internal forces and torques related to residual stresses of the II. type are in an equilibrium across a sufficient number of grains. When interfering with this equilibrium, macroscopic dimension changes may develop.
x 0
grain boundaries
Residual stresses of the III. type are inhomogenuous across smallest material areas (some atomic distances). Internal forces and torques related to residual stresses of the III. type are in an equilibrium across small areas (sufficiently large part of a grain). When interfering with this equilibrium, macroscopic dimension changes do not develop.
sE = s I + sII
sIII
<
= residual stresses between several grains = residual stresses in a single grain
< <
sI sII sIII
+
= residual stresses in a point
© ISF 2002
br-er03-03e.cdr
© ISF 2002
br-er03-04e.cdr
Definition of Residual Stresses
Definition of Residual Stresses of I., II., and III. Type
Figure 3.3
Figure 3.4
homogeneous. The left part of 1000 °C 900
Figure 3.5 shows the T-t°C
urement points in the cylinder. of quenching on the stress condition in the cylinder. At
Temperature
Figure 3.6 shows the results
1
750
2
3
35 mm diameter water cooling 500
250
MS
1 edge 2 50 % radius 3 core
1s
5s
15 s
800
1000
Temperature
curve of three different meas-
0s 10 s
700
20 s
600
25 s
500
35 s 400
45 s
300
53 s 200
the beginning of cooling, the
0 -2 10
68 s 10-1
cylinder edge starts shrinking
10-0
101 102 Cooling time
103
s 104
100 280 s
0 17,5
7
14 10,5
faster than the core (upper
7
0 3,5
3,5
10,5
Radius © ISF 2002
br-eI-03-05e.cdr
figure). Through the stabilising
Temperature in a Cylinder During Water Cooling
effect of the cylinder core, Figure 3.5
mm 17,5
3. Residual Stresses
24
tensile stress builds up at the edge areas while the core is exposed to pressure stress. Resulting volume differences between core and edge are balanced by elastic and plastic deformations. When cooling is completed, edge and core are on the same temperature level, the plastically stretched edge now supports the unstressed core, so that pressurestresses are present in the edge areas and tensile residual stresses in the core.
300
tension pressure
N/mm²
E
200
tension
Stresses in the central rod
Volume differences between edge and core at start of cooling
tension pressure
tension
Compensation of volume differences by plastic deformation and stresses at start of cooling
pressure
D
100
0
A
C
-100
-200
B'
tension
B
pressure
br-er03-06e.cdr
Compensation of volume differences by plastic deformation and stresses at end of cooling
-300 0 © ISF 2002
400
°C
600
br-er03-07e.cdr
© ISF 2002
Residual Stress Development by Warming the Central Rod
Volume Changes During Cooling
Figure 3.6
200
Temperature of the central rod
Figure 3.7
These changes are principally shown once again in Figure 3.7 with the 3-rod model. A warming of the middle rod causes at first an elastic expansion of the outer rods, the inner rod is exposed to pressure stress (line A-B). Along the line B-C the rod is plastically deformed, because pressure stresses have exceeded the yielding point. At point C, the cooling of the rod starts, it is exposed to tensile stress due to shrinking. Along the line D-E the rod is plastically deformed due to the influence of the counter members beeing in tension. At the point E the system has cooled down to its initial temperature. This point represents the remaining residual stress condition of this construction. If heating is stopped before point C is reached and cooled down to the initial temperature, then stress increase in the centre rod will be in parallel
3. Residual Stresses
25
with the elastic areas. Starting with point B, the same residual stress condition is present as in a case of heating up to a temperature above 600°C. Figure 3.8 divides the development of residual stresses in welded seams in three different mechanisms. Shrinking stresses: these are stresses formed through uniform cooling of the seam. Caused by expansion restriction of the colder areas at the edge of the weld and base material , tensile stresses develop along and crosswise to the seam. Quenching stresses: If cooling is not homogenous, the surface of the weld cools down faster than the core areas. If the high-temperature limit of elasticity is exceeded due to buildup stress differences, pressure stresses will be present at the weld surface after cooling. In contrast, the core shows tensile stresses in cold condition (see also Figure 3.6). Transition stresses: Transitions in the ferrite and perlite stage cause normally only residual stresses, because within this temperature range the yield strength of the steel is so low that generated stresses can be undone by plastic deformations. This is not the case with transitions in the Bainite and martensite stage. A transition of the austenite causes an increase in volume (transition cfc in cbc, the cfc lattice has a higher density, additional volume increase through lat+y
tice deformation). In the case of a homoge-
-x
nous transition, the weld will consequently unfold pressure stresses. If the transition of +x
the edge areas happens earlier than the transition of the slower cooling core, plastic de-
-y 2. Quenching stresses
1. Shrinking stresses
formations of the core area may be present similar to quenching (see above: quenching
-x
+x
-x
+s +y
-s -y
+x
stresses). In this case, the weld surface will 3. Transformation stresses
show tensile stresses after cooling. Generally these mechanisms cannot be separated accurately from each other, thus
4. Overlap options of case 1., 2. and 3.
+s +y
+s +y
inhomogenuous transformation
-x
+x
-x
+x
the residual stress condition of a weld will represent an overlap of the cases as shown in the 3rd figure. This overlap of the different
homogenuous transformation
-s -y br-er03-08.cdr
-s -y © ISF 2002
Stress Distributions and Superpositions Perpendicular to Welded Joint
mechanisms makes a forecast of the remaining residual stress condition difficult. Figure 3.8
3. Residual Stresses
26
Figure 3.9 shows the building-up of residual Temperature distribution
Seam
Stress distribution sX
ogy to the 3-rod model of Figure 3.7. This fig-
1. cut A-A DT ~ 0
x
stresses crosswise to a welded seam in anal-
stress-free
ure considers only shrinking residual stresses. Before application of welding heat, the seam
A
A
2. cutt B-B
area is stress-free (cut A-A). At the weldpool tension
weldpool B
the highest temperature of the welding cycle
B
area of plastic deformations
C
pressure
C
can be found (cut B-B), metal is liquid. At this point, there are no residual stresses, because
3. cut C-C
molten metal cannot transmit forces at the D
D
M
weldpool. Areas close to the joint expand through welding heat but are supported by
M'
4. cut D-D
residual stresses
areas which are not so close to the seam.
DT = 0
Thus, areas close to the joint show compres© ISF 2002
br-er03-09e.cdr
Formation of Residual Stresses Caused by Welding Heat
sion stress, areas away from the joint tensile stress. In cut C-C the already solidified weld metal starts to shrink and is supported by
Figure 3.9 areas close to the seam, the weld metal shows tensile stresses, the adjacent areas compression stresses. In cut D-D is the temperature completely balanced, a residual stress condition is recognised as shown in the lower right figure. 31 15 mm 15 mm
material S235JR (St 37)
103 a a
Figure 3.10 shows how much residual stresses are influenced by constraining ef-
1.
a = 100 mm
s = 800 N/mm²
fects of adjacent material. The resulting
2.
a = 150 mm
s = 530 N/mm²
stress in the presented case is calculated
3.
a = 200 mm
s = 400 N/mm²
according to Hooke:
4.
a = 250 mm
s = 300 N/mm²
σ= ε·E
5.
a = 300 mm
s = 270 N/mm²
br-er03-10e.cdr
© ISF 2002
Shrinking Stresses in a Firmly Clamped Plate
Elongation ε is calculated as ∆ l/a (∆ l is the length change due to shrinking). With conFigure 3.10
3. Residual Stresses
27
stant joint volume will shrinking and ∆ l always have the same value. Thus the elongation ε depends only on the value a. The smaller the a is chosen, the higher are the resulting stresses. Effects of transition on cooling can be estimated from Figure 3.11. Here curves of temperature- and length-changes of ferritic and austenitic steels are drawn. It is clear that a ferritic lattice has a higher volume than an austenitic lattice at the same temperature. A steel which transforms from austenite to one of the ferrite types increases its volume at the critical point. This sudden rise in volume can be up to 3% in the case of martensite formation.
Longitudinal expansion Dl
welding sample 300 x 10 x 30 (70,140) groove angle 60°, depth 4,5 mm
firm clamping
force sensor
el el
thermo couples
links
ste
nit ic
ste
tic
rri
fe
ste
au
to calculator
1000
N
°C
600
800
14
m tra ild ns ste fo el rm w at ith io n
800
200
Temperature
Force
elektrode 400
600
heat affected zone
400
force 0
Temperature [°C]
200
temperature -200
0 -1 10
100
101
102
103
104
s
105
Time br-er03-12e.cdr br-er03-11e.cdr
© ISF 2002
© ISF 2002
Force Measurement During Cooling of a Weld
Longitudinal Expansion of Various Steels
Figure 3.11
Figure 3.12
To record the effects of this behaviour on the stress condition of the weld, sample welds are carried out in the test device outlined in Figure 3.12. Thermo couples measure the T-t – curve at the weld seam, a force sensor records the force which tries to bend the samples. The lower picture shows the results of such a test. The temperature behaviour at the fusionline as well as the force necessary to hold the sample over the time is plotted.
3. Residual Stresses
28
In the temperature range above 600°C the force sensor registers a tensile force which is caused by the shrinking of the austenite. Between 600 and 400°C a large drop in force can be seen, which is caused by the transition of the austenite. The repeated increase of the force is based on further shrinking of the ferrite. With the help of TTT diagrams of base material and welding
steel
austenitic
S690QL (StE 70)
consumable,
consumable electrode
austenitic
austenitic
surface weld
surface weld
the
transition
temperatures and/or temperature areas for the individual zones of the welded joint can
S690QL (StE 70) high-strength
sample shape (V-groove, 60°) type of welding
surface weld
position of the HAZ
data and with the course of residual stress distribution sL
0
pressure
temperature it can be clearly
tension
be determined. With these
determined in which part of
© ISF 2002
br-eI-03-13e.cdr
the curve the force drop is
Influence of Material Combination on Residual Stress Distribution in a Weld
caused by the transition of the welding consumable and in
Figure 3.13 which part by transition in the heat affected 5°42'
2°8'
1°51'
zone (HAZ). These results can be used to determine the longitudinal residual stresses transversal to the joint, as shown in Figure 3.13. During
140
welding of austenitic transition-free materials
Angle change
% 100
only tensile residual stresses are caused in
80 60
the welded area according to Figure 3.8. If an
40 20
austenitic electrode is welded to a StE 70, transitions occur in the area of the heat af-
f = 1°
f = 3°
f = 7°
fected zone which lead to a decrease of ten-
f = 13°
sile stresses. If a high-strength electrode which has a martensitic transition, is welded a=5
a=7
a=9
br-er03-14e.cdr
© ISF 2002
Influence of Welding Sequence on Angle Distortion
Figure 3.14
a = 12,5
to a StE 70, then there will be pressure residual stresses in the weld metal and tensile residual stresses in the HAZ.
3. Residual Stresses
29
If parts to be welded are not fixed, the shrinking of the weld will cause an angular distortion of the workpieces, Figure 3.14 . If the workpieces can shrink unrestricted in this way, the remaining residual stresses will be much lower than in case with firm clamping. Methods to determine residual stresses can be divided into
destructive,
plan
section
nona
destructive, and conditionWSG
ally destructive methods. The borehole and ring core
c
method can be considered b
as conditionally destructive, workpiece
Figures 3.15 and 3.16. In both cases, present re-
© ISF 2002
br-eI-03-15e.cdr
sidual stresses are released
Residual Stress Determination Using Bore Hole Procedure
through partial material removal and the resulting deformations
are
Figure 3.15
then
measured by wire strain gauges. An essential advantage of the borehole method is the very small material removal, the diameter of the borehole is only 1 to 5 mm, the bore depth is 1- to 2-times the borehole diameter. The disadvantage here is that only surface elongations can be measured, thus the results are limited residual stresses in the surface area of the workpiece. wire expansion gauge
With the ring core method,
b(sb) 45°
c(ec)
a crown milling cutter is z
45°
t0.21
a(sa)
D1.58
used to mill a ring groove around a three-axes wire strain gauge. The core is
s1 (z)
s2 (z)
measurement point
released from the force effects and stress-relieved. At the time when the resil-
© ISF 2002
br-eI-03-16e.cdr
Residual Stress Determination Using Ring Core Procedure
Figure 3.16
ience of the core is measured, the detection of the residual stress distribution
3. Residual Stresses
30
across the depth is also possible. Both methods are limited in their suitability for measuring welding residual stresses, because steep strain gradients in the HAZ may cause wrong measurements.
destructive
measurement
A
A
A
mechanical deformations
A E
A
A E
A E
A E
A E
optical procedures
A E
magnetic
thermal processes
ultra sonic
breaking-up bending deflection
ring groove
methods.
cam web
ing one of the respective
drilling out turning off
to be picked-up when us-
causes
others
x - ray
mechanically - electrically
sidual stresses and what causes residual stresses
partial
optical procedures
urement methods for re-
complete
bore hole
shows a survey of meas-
non-destructive
ring core
The table in Figure 3.17
A E
A E
E
E
A E
A E
E
E
A E
A E
E
E
surfacetreatment
A - general application E - further development desired © ISF 2002
br-eI-03-17e.cdr
Methods for Determination of Residual Stresses
Figure 3.17
assumption of stress distribution
Figure 3.18 shows a sur-
measured variable
residual stresses
cutting in layers
vey of the completely destructive
procedures
f
biaxial
f
any
y
0
of
bending deflection f curves reduced curves
sy sz tzy
tear f
partial residual stress relief by Dsz
z
x
cutting-in
residual stress recognition.
f
uniaxial locally different linear, tensile residual stresses on top, down pressure stresses
drilling e45 eT eL
slitting 0.46f
tripleaxial independent of smple length sL, sT, sR
uniaxial linear symmetrically with reference to rod axis
length change eL circumference change eT
tear f
sL sT sR
partial residual stress relief by Dsz
© ISF 2002
br-eI-03-18e.cdr
Destructive Methods for Determination of Residual Stresses
Figure 3.18
4. Heat Treatment and its Function During Welding
4. Heat Treatment and its Function During Welding
32
When welding a workpiece, not only the weld itself, but also the surrounding base material (HAZ) is influenced by the supplied heat quantity. The temperature-field, which appears around the weld when different welding procedures are used, is shown in Figure 4.1.
Figure 4.2 shows the influence of the material properties on the welding process. The determining factors on the process presented in this Figure, like melting temperature and interval, heat capacity, heat extension etc, depend greatly on the chemical composition of the material. Metallurgical properties are here characterized by e.g. homogeneity, structure and texture, physical properties like heat extension, shear strength, ductility. Figure 4.1
Structural changes, caused by the heat input
(process 1, 2, 7, and 8), influence directly the mechanical properties of the weld. In addition, the chemical composition of the weld metal and adjacent base material are also influenced by the processes 3 to 6.
Based on the binary system, the formation of the different structure zones is shown in Figure 4.3. So the coarse grain zone occurs in areas of intensely
elevated
austenitising temperature for example. At the same time, hardness peaks appear in these
areas
because
of
greatly reduced critical cooling rate and the coarse austenite Figure 4.2
4. Heat Treatment and its Function During Welding
33
grains. This zone of the weld is the area, where the worst toughness values are found.
In Figure 4.4 you can see how much the formation of the individual structure zones and the zones of unfavourable mechanical properties can be influenced. Applying an electroslag one pass weld of a 200 mm thick plate, a HAZ of approximately 30 mm width is achieved. Using a three pass technique, the HAZ is reduced to only 8 mm.
With the use of different procedures, the differences in the formation of heat affected zones become even clearer as shown in Figure 4.5. These effects can actively be used to the advantage of the material, for example to adjust
Figure 4.3
calculated mechanical properties to one's choice or to remove negative effects of a welding. Particularly with high-strength fine grained steels and high-alloyed materials, which are specifically optimised to achieve special quality, e.g. corrosion resistance against a certain attacking medium, this post-weld heat treatment is of great importance.
Figure 4.6 shows areas in the Fe-C diagram of different heat treatment methods. It is clearly visible that the carbon content (and also the content of other alloying elements) has a distinct influence on the level of annealing temperatures like e.g. coarse-grain
Figure 4.4
4. Heat Treatment and its Function During Welding
34
heat treatment or normalising. It can also be seen that the start of martensite formation (MS-line) is shifted to continuously
Figure 4.5
Figure 4.6
decreasing temperatures with increasing C-content. This is important e.g. fo r hardening processes (to be e xplained later).
As this diagram does not cover the time influence, only
constant
stop-
temperatures can be read, predictions about heating-up and cooling-down rates are not possible. Thus the individual heat treatment methods will be explained by their
temperature-time-
behaviour in the following. Figure 4.7
4. Heat Treatment and its Function During Welding
35
Figure 4.7 shows in the detail to the right a T-t course of coarse grain heat treatment of an alloy containing 0,4 % C. A coarse grain heat treatment is applied to create a grain size as large as possible to improve machining properties. In the case of welding, a coarse grain is unwelcome, although unavoidable as a consequence of the welding cycle. You can learn from Figure 4.7 that there are two methods of coarse grain heat treatment. The first way is to austenite at a temperature close above A3 for a couple of hours followed by a slow cooling process. The second method is very important to the welding process. Here a coarse grain is formed at a temperature far above A 3 with relatively short periods. Figure 4.8 shows schematically time-temperature behaviour
in
a
TTT-
diagram. (Note: the curves explain running structure mechanisms, they must not be used as reading off examples. To determine t8/5, hardness values, or microstructure distribution, are TTT-diagrams always read continuously Figure 4.8
mally.
Mixed
or
isother-
types
like
curves 3 to 6 are not a llowed for this purpose!).
The most important heat treatment methods can be divided into sections of annealing, hardening and tempering, and these single processes can be used individually or combined. The normalising process is shown in Figure 4.9. It is used to achieve a homogeneous ferrite perlite structure. For this purpose, the steel is heat treated approximately 30°C above Ac3 until homogeneous auste nite evolves. This condition is the starting point for the following hardening and/or quenching and tempering treatment. In the case of hypereutectoid steels, austenisation takes place above the A1 temperature. Heating-up should be fast to keep the austenite grain as fine as possible (see TTA-diagram, chapter 2). Then air cooling follows, leading normally to a transformation in the ferrite condition (see Figure 4.8, line 1; formation of ferrite and perlite, normalised micro-structure).
4. Heat Treatment and its Function During Welding
36 To harden a material, austenisation and homogenisation is carried out also at 30°C above AC3. Also in this case one must watch that the austenite grains remain as small as possible. To ensure a complete transformation to marte nsite, a subsequent quenc hing
follows
until
the
temperature is far below Figure 4.9
the Ms-temperature, Figure 4.10. The cooling rate dur-
ing quenching must be high enough to cool down from the auste nite zone directly into the martensite zone without any further phase transitions (curve 2 in Figure 4.8). Such quenching processes build-up very high thermal stresses which may destroy the workpiece during hardening. Thus there are variations of this process, where perlite formation is suppressed, but due to a smaller temperature gradient thermal stresses remain on an uncritical level (curves 3 and 4 in Figure 4.8). This can be achieved in practice –for example- through stopping
a
water
quenc hing
process at a certain temperature and continuing the cooling with a milder cooling medium (oil). With longer holding on at elevated temperature level, transformations can also be carried through in the bainite area (curves 5 and 6).
Figure 4.10
4. Heat Treatment and its Function During Welding
37
Figure 4.11 shows the quenching and tempering procedure. A hardening is followed by another heat treatment below Ac1. During this tempering process, a break down of °C
austenite
about 30°C above A3
900
A1
700 ferrite + perlite
takes
place.
Ferrite and cementite are
A3
formed. As this change
Temperature
austenite + ferrite Temperature
martensite
hardening and tempering
quenching
causes a very fine micro-
slow cooling
500
structure, this heat treatment leads to very good
300 0,4
0,8 C-Content
%
mechanical properties like
Time
br-eI-04-11.cdr
e.g. strength and toughHardening and Tempering
ness.
Figure 4.11 Figure 4.12 shows the procedure of soft-annealing. Here we aim to adjust a soft and suitable micro-structure for machining. Such a structure is characterised by mostly globular formed cementite particles, while the lamellar structure of the perlite is resolved (in Figure 4.12 marked by the circles, to the left: before, to the right: after soft-annealing). For hypoeutectic steels, this spheroidizing of cementite is achieved by a heat treatment close below A1. With these steels, a part of the cementite bonded carbon dissolves during heat treating close below A1, the remaining cementite lamellas transform with time into balls, and the bigger ones grow at the expense of the smaller ones (a transfor-
°C
time dependent on workpiece
mation is carried out be-
modynamically more favourable
austenite + ferrite
oscillation annealing + / - 20 degrees around A 1
10 to 20°C below A1
A3 A1
Temperature
strongly reduced → ther-
900
Temperature
cause the surface area is
austenite
700 ferrite + perlite
or
500
condition). 300
Hypereutectic steels have
0,4
in addition to the lamellar
0,8 C-Content
%
Time cementite
structure of the perlite a br-eI-04-12.cdr
cementite network on the
Soft Annealing
grain boundaries.
Figure 4.12
4. Heat Treatment and its Function During Welding
38
Spheroidizing of cementite is achieved by making use of the transformation processes during oscillating around A1. When exceeding A1 a transformation of ferrite to auste nite takes place with a simultaneous solution of a certain amount of carbon according to the binary system Fe C. When the temperature drops below A1 again and is kept about 20°C below until the transformation is completed, a re-precipitation of cementite on existing nuclei takes place. The repetition of this process leads to a stepwise
spheroidizing
of
cementite and the frequent transformation
avoids
a
grain coarsening. A softannealed
microstructure
represents frequently the delivery condition of a material.
Figure 4.13
Figure 4.13 shows the principle of a stress-relieve heat treatment. This heat treatment is used to eliminate dislocations which were caused by welding, deforming, transformation etc. to improve the toughness of a workpiece. Stress-relieving works only if present dislocations are able to move, i.e. plastic structure deformations must be executable in the micro-range. A temperature the
increase
commonly
is
used
method to make such deformations
possible
be-
cause the yield strength limit decreases with increasing temperature. A stress-relieve heat treatment should not cause any other change to properties, so that tempering steels
Figure 4.14
4. Heat Treatment and its Function During Welding
39
are heat treated below tempering temperature. Figure 4.14 shows a survey of heat treatments which are important to welding as well as their purposes.
Figure 4.15 shows principally the heat treatments in connection
with
welding.
Heat treatment processes are divided into: before, during, and after welding. Normally a stress-relieving or normalizing heat treatment
is
applied
before
welding to adjust a proper material condition which for welding. After welding, alFigure 4.15 most any possible heat treatment can be carried out. This is only limited by workpiece dimensions/shapes or arising costs. The most important section of the diagram is the kind of heat treatment which accom-panies the welding. The most important processes are e xplained in the follo wing.
Figure 4.16 represents the influence of different accompanying heat treatments during welding, given within a TTT-diagram. The fastest cooling is achieved with welding without preheating, with addition of a small share of bainite, mainly martensite is formed (curve 1, analogous to Figure 4.8, hardening). A simple heating before welding without additional stopping time lowers the cooling rate according to curve 2. The proportion of martensite is reduced in the forming structure, as well as the Figure 4.16
4. Heat Treatment and its Function During Welding
40
level of hardening. If the material is hold at a temperature above MS during welding (curve 3), then the martensite formation will be completely suppressed (see Figure 4.8, curve 4 and 5).
To explain the temperature-time-behaviours used in the following, Figure 4.17 shows a superposition of all individual influences on the materials as well as the resulting T-Tcourse in the HAZ. As an example, welding with simple preheating is selected. The plate is preheated in a period tV . After removal of the heat source, the cooling of the workpiece starts. When t S is reached, welding starts, and its temperature peak overlays the cooling curve of the base material. When the welding is completed, cooling period tA starts. The full line represents the resulting temperature-time-behaviour of the HAZ.
The temperature time course during welding with simple preheating is shown in Figure
Figure 4.17 4.18. During a welding time tS a drop of the working temperature TA occurs. A further air cooling is usually carried out, however, the cooling rate can also be reduced by cove ring with heat insulating materials.
Another variant of welding with preheating is welding at Figure 4.18
constant
temperature.
working This
is
4. Heat Treatment and its Function During Welding
41 achieved through further warming during welding to avoid a drop of the working temperature. In Figure 4.19 is this case (dashed line, TA needs not to be above MS) as well as the special case of isothermal welding illustrated. During isothermal welding, the workpiece is heated up to a working temperature
Figure 4.19
above
MS
(start of martensite formation) and is also held there
after welding until a transformation of the austenitised areas has been completed. The aim of isothermal welding is to cool down in accordance with curve 3 in Figure 4.16 and in this way, to suppress martensite formation.
Figure 4.20 shows the T-T course during welding with post-warming (subsequent heat treatment, see Figure 4.15). Such a treatment can be carried out very easy, a gas welding torch is normally used for a local preheating. In this way, the toughness properties of some steels can be greatly improved. The lower sketch shows a combination of pre- and postheat treatment. Such a treatment is applied to steels which have such a strong tendency to hardening that a cracking in spite of a simple preheating before welding cannot be avoided, if they cool down directly from working temperature. Such materials are heat treated immediately after welding at a temperature between 600 and 700°C, so that a formation Figure 4.20
4. Heat Treatment and its Function During Welding
42
of martensite is avoided and welding residual stresses are eliminated simultaneously.
Aims of the modified stephardening
welding
should
not be discussed here, Figure 4.21. Such treatments are used for transformationinert materials. The aim of the figure is to show how complicated a heat treatment can become for a material in combination with welding.
Figure 4.22 shows temperature distribution during multi-
Figure 4.21
pass welding. The solid line represents the T-T course of a point in the HAZ in the first pass. The root pass was welded without preheating. Subsequent passes were welded without cooling down to a certain temperature. As a result, working temperature increases with the number of passes. The second pass is welded under a preheat temperature which is already above martensite start temperature. The heat which remains in the workpiece preheats the upper layers of the weld, the root pass is post-heat treated through the same effect. During welding of the last pass, the preheat temperature has reached such a high level that the critical cooling rate will not be surpassed. A fa vourable effect of multi-pass welding is the warming of the HAZ of each previous pass above recrystallisation temperature with the corresponding crystallisa-
Figure 4.22
4. Heat Treatment and its Function During Welding
43
tion effects in the HAZ. The coarse grain zone with its unfavourable mechanical properties is only present in the HAZ of the last layer. To achieve optimum mechanical values, welding is not carried out to Figure 4.22. As a rule, the same welding conditions should be applied for all passes and prescribed t8/5 – times must be kept, welding of the next pass will not be carried out before the previous pass has cooled down to a certain temperature (keeping the interpass temperature). In addition, the workpiece will not heat up to excessively high temperatures.
Figure 4.23 shows a nomogram where working te mperature and minimum and maximum heat input for some steels can be interpreted, depending on carbon equivalent and wall thickness. If e.g. the water quenched and tempered fine grain structural steel S690QL of 40 mm wall thickness is welded, the following data can be found:
- minimum heat input between 5.5 and 6 kJ/cm - maximum heat input about 22 kJ/cm - preheating to about 160°C - after welding, residual stress relieving between 530 and 600°C.
Steels which are placed in the
hatched
soaking
area,
area
called
must
be
treated with a hydrogen relieve annealing. Above this area, a stress relieve annealing must be carried out. Below this area, a post-weld heat treatment is not required.
Figure 4.23
5. Welding Plain and Low Alloy Steels
5. Welding Plain and Low Alloy Steels
45
tiD o e f n
g n u e lti E
e d fs n iB rg
ra h m e c d tlS ä
a n d e r
ta le h S s
h Z s a ic n m e u
ä tilh e rg S
cn rh i t o
s n m e g tz u
le
e tS slä n d h
a n d e r
ä tilh e rg S
le -rg ti
h lu tä Q a si
e
o g n u ltce irh E
e ä S sld n r tü p g h u a c H
u -g n ile rs k a
sä tu e Q a lid E h
e l-rg it
b
r-e 0 d 5 c1 .
e h l
lh sd E e ä t
2 0 S F I©
4
In the European Standard DIN EN
10020 (July 2000), the designations
Definition of the term “steel” Steel is a material with a mass fraction if iron which is higher than of every other element, ist carbon content is, in general, lower than 2% and steel contains, moreover, also other elements. A limited number of chromium steels might contain a carbon content which is higher than 2%, but, however, 2% is the common boundary between steel and cast iron [DIN EN 10020 (07.00)].
(main symbols) for the classification of steels are standardised. Figure 5.1 shows the definition of the term „steel“ and the classification of the steel
Classification in accordance with the chemical composition: l
unalloyed steels
l
stainless steels
l
other, alloyed steels
grades
in
accordance
with
their
chemical composition and the main quality classes.
Classification in accordance with the main quality class: · unalloyed steels
- unalloyed quality steels - unalloyed special steels
· stainless steels · other, alloyed steels
- alloyed quality steels - alloyed special steels © ISF 2004
br-er05-01.cdr
Definition for the classification of steels
Figure 5.1 In accordance with the chemical compoDetermined element
sition the steel grades are classified into Al
aluminium
unalloyed, stainless and other alloyed
B
boron
Bi
bismuth
steels. The mass fractions of the individ-
Co cobalt
ual elements in unalloyed steels do not
Cu copper
achieve the limit values which are indicated in Figure 5.2. Stainless steels are grades of steel with a mass fraction of chromium of at least 10,5 % and a maximum of 1,2 % of carbon. Other alloyed steels are steel grades which do not comply with the definition of stainless steels and where one alloying
Cr
limit value Mass fraction in %
chromium
La
lanthanides (rated individually) Mn manganese Mo molybdenum Nb niobium Ni
nickel
Pb lead Se selenium Si
silicon
Te
tellurium
Ti V
titanium vanadium
W
tungsten
Zr zirconium Others (with the exception of carbon, phosphorus, sulphur, nitrogen) (Each) a) If just the highest value has been determined for mangenese, the limit value us 1,80% and the 70%-rule does not apply. br-er05-02.cdr
element exceeds the limit value indicated
© ISF 2004
Boundary between unalloyed and alloyed steels
in Figure 5.2. Figure 5.2
5. Welding Plain and Low Alloy Steels
46
As far as the main quality classes are concerned, the steels are classified in accordance with their main characteristics and main application properties into unalloyed, stainless and other alloyed steels. As regards unalloyed steels a distinction is made between unalloyed quality steels and unalloyed high-grade steels. Regarding unalloyed quality steels, prevailing demands apply, for example, to the toughness, the grain size and/or the forming properties. Unalloyed high-grade steels are characterised by a higher degree of purity than unalloyed quality steels, particularly with regard to non-metal inclusions. A more precise setting of the chemical composition and special diligence during the manufacturing and monitoring process guarantee better properties. In most cases these steels are intended for tempering and surface hardening. Stainless steels have a chromium mass fraction of at least 10,5 % and maximally 1,2 % of carbon. They are further classified in accordance with the nickel content and the main characteristics: corrosion resistance, heat resistance and creep resistance. Other alloyed steels are classified into alloyed quality steels and alloyed high-grade steels. Special demands are put on the alloyed quality steels, as, for example, to toughness, grain size and/or forming properties. Those steels are generally not intended for tempering or surface hardening. The alloyed high-grade steels comprise steel grades which have improved properties through precise setting of their chemical composition and also through special manufacturing and control conditions.
5. Welding Plain and Low Alloy Steels
47
The European Standard DIN EN 10027-1 (September 1992) stipulates the rules for the designation of the steels by means of code letters and identification numbers. The code letters and identification numbers give information about the main application field, about the mechanical or physical properties or about the composition. The code designations of the steels are divided into two groups. The code designations of the first group refer to the application and to the mechanical or physical properties of the steels. The code designations of the second group refer to the chemical composition of the steels. l
S = Steels for structural steel engineering e.g. S235JR, S355J0
According to the utilization of the
l
P = Steels for pressure vessel construction e.g. P265GH, P355M
steel and also to the mechanical or
l
L = Steels for pipeline construction e.g. L360A, L360QB
physical properties, the steel grades
l
E = Engineering steels e.g. E295, E360
l
B = Reinforcing steels e.g. B500A, B500B
l
Y = Prestressing steels e.g. Y1770C, Y1230H
l
R = Steels for rails (or formed as rails) e.g. R350GHT
l
H = Cold rolled flat-rolled steels with higher-strength drawing quality e.g. H400LA
l
D = Flat products made of soft steels for cold reforming e.g. DD14, DC04
l
T = Black plate and tin plate and strips and also specially chromium-plated plate and strip e.g. TH550, TS550
l
M = Magnetic steel sheet and strip e.g. M400-50A, M660-50D
br-er05-03.cdr
of the first group are designated with different main symbols (Fig. 5.3).
© ISF 2004
Classification of steels in accordance with their designated use
Figure 5.3
5. Welding Plain and Low Alloy Steels
48
An example of the code designation structure with reference to the usage and the mechanical or physical properties for “steels in structural steel engineering“ is explained in Figure 5.4.
Figure 5.4
5. Welding Plain and Low Alloy Steels
49
For designating special features of the steel or the steel product, additional symbols are added to the code designation. A distinction is made between symbols for special demands, symbols for the type of coating and symbols for the treatment condition. These additional symbols are stipulated in the ECISS-note IC 10 and depicted in Figures 5.5 and 5.6.
Symbol1)2)
Coating
+A + AR + AS + AZ + CE + Cu + IC + OC +S + SE +T + TE +Z + ZA + ZE + ZF + ZN
hot dipped aluminium, cladded by rolling coated with Al-Si alloy coated with Al-Tn alloy (>50% Al) electrolytically chromium-plated copper-coated inorganically coated organically coated hot-galvanised electrolytically galvanised upgraded by hot dipping with a lead-tin alloy electrolytically coated with a lead-tin alloy hot-galvised coated with Al-Zn alloy (>50% Zn) electrolytically galvanised diffusion-annealed zinc coatings (galvannealed, with diffused Fe) nickel-zinc coating (electrolytically) 1 2
) The symbols are separated from the preceding symbols by plus-signs (+) ) In order to avoid mix-ups with other symbols, the figure S may precede,
for example +SA © ISF 2004
br-er-05-05.cdr
Symbols for the coating type
Figure 5.5
1 2
Symbol ) )
treatment condition
+A + AC +C
softened annealed for the production of globular carbides work-hardened (e.g., by rolling and drawing), also a distinguishing mark for cold-rolled narrow strips) cold-rolled to a minimum tensile strength of nnn MPa/mm² cold-rolled thermoformed/cold formed slightly cold-drawn or slightly rerolled (skin passed) quenched or hardened treatment for capacity for cold shearing solution annealed untreated
+ Cnnn + CR + HC + LC +Q +S + ST +U
1 2
) The symbols are separated from the preceding symbols by plus-signs (+) ) In order to avoid mix-ups with other symbols, the figure T may precede,
for example +TA © ISF 2004
br-er-05-06.cdr
Symbols for the treatment condition
Figure 5.6
5. Welding Plain and Low Alloy Steels
50
Figure 5.7 shows an example of the novel designation of a steel for structural steel engineering which had formerly been labelled St37-2.
The steel St37-2 (DIN 17100) is, according to the new standard (DIN EN 10027-1), designated as follows:
S235 J 2 G3 further property (RR = normalised)
Steel for structural steel engineering
ReH ³ 235 MPa/mm2
test temperature 20°C impact energy ³ 27 J
S = steels for structural steel engineering P = steels for pressure vessel construction L = steels for pipeline construction E = engineering steels B = reinforcing steels © ISF 2002
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Steel designation in accordance with DIN EN 10027-1
Figure 5.7 Steel Stahl S355J0 (St 52-3) S500N (StE500) P295NH (HIV) S355J2G1W (WTSt510-3) S355G3S (EH36)
C
Si
Mn
P
S
Cr
Al
Cu
N
Mo
Ni
Nb
V
£0,20
£0,55
£1,60
0,040
0,040
/
/
/
£0,009
/
/
/
/
0,1 - 0,6 1 - 1,7
0,035
0,030
0,30
0,020
0,20
0,020
0,1
1
0,05
0,22
0,21 £0,26
£0,35
£0,05
£ 0,05
/
/
/
/
/
/
/
/
£0,15
£0,50 0,5 - 1,3 0,035
0,035
0,40 0,80
/
0,25 0,5
/
£0,30
£0,65
/
0,02 0,12
£ 0,18
£0,1 0,7 - 1,5 £0,05 0,35
£ 0,05
/
/
/
/
/
/
/
/
Steel Stahl
³0,6
Tensile strength Zugfestigkeit RmRm [N/mm²]
yield point ReeHH Streckgrenze [N/mm²]
elongation after fracture Bruchdehnung A A [%]
impact energy AVV Kerbschlagarbeit [J] -20°C
0°C S355J2G3 (St 52-3) S500N (StE500) P295NH (HIV) S355J2G1W (WTSt510-3) S355G3S (EH36)
510-680
355
20-22
27 31-47
610-780
500
16
460-550
285
>18
510-610
355
22
400-490
355
>22
27 21-39
49 (bei +20°C)
76 (bei -10°C) © ISF 2004
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Chemical composition and mechanical parameters of different steel sorts
Figure 5.8 Figure 5.8 depicts the chemical composition and the mechanical parameters of different steel grades. The figure explains the influence of the chemical composition on the mechanical properties.
5. Welding Plain and Low Alloy Steels
51
The steel S355J2G2 represents the basic type of structural steels which are nowadays commonly used. Apart from a slightly increased Si content for desoxidisation it this an unalloyed steel. S500N is a typical fine-grained structural steel. A very fine-grained microstructure with improved tensile strength values is provided by the addition of carbide forming elements like Cr and Mo as well as by grain-refining elements like Nb and V. The boiler steel P295NH is a heat-resistant steel which is applied up to a temperature of 400°C. This steel shows a relatively low strength but very good toughness values which are caused by the increased Mn content of 0,6%. S355J2G1W is a weather-resistant structural steel with mechanical properties similar to S355J2G2. By adding Cr, Cu and Ni, formed oxide layers stick firmly to the workpiece surface. This oxide layer prevents further corrosion of the steel. S355G3S belongs to the group of shipbuilding steels with properties similar to those of usual structural steels. Due to special quality requirements of the classification companies (in this case: impact energy) these steels are summarised under a special group.
5. Welding Plain and Low Alloy Steels
52
The steel grades are classified into four subgroups according to the chemical composition (Fig 5.9): ● Unalloyed steels (except free-cutting steels) with a Mn content of < 1 % ● Unalloyed steels with a medium Mn content > 1 %, unalloyed free-cutting steels and alloyed steels (except high-speed steels) with individual alloying element contents of less than 5 percent in weight ● Alloyed steels (except high-speed steels), if, at least for one alloying element the content is ≥ 5 percent in weight ● High-speed steels
The unalloyed steels with Mn conUnalloyed steels (Mo content < 1%)
tents of < 1% are labelled with the
C45
code letter C and a number which
0,45% Carbon
Carbon
complies with the hundredfold of the
Unalloyed steels (Mn content > 1%)
10CrMo9-10
mean value which is stipulated for the carbon content.
C=10/100=0,10%
Cr=9/4=2,25% element
Unalloyed steels with a medium Mn
Mo=10/10=1% factor
Cr, Co, Mn, Ni, Si, W
4
Al, Be, Cu, Mo, Nb, Pb, Ta, Ti, V, Zr
content > 1 % are labelled with a
C, Ce, N, P, S
10 100 1000
B
Table 5.1
number which also complies with a
Alloyed steels (content of alloying element > 5%)
hundredfold of the mean value which
X10CrNi18-10
is stipulated for the carbon content, the
Legiert
C=10/100=0,1%
chemical symbols for the alloying
Cr=18%
Ni=10%
High-speed steels
elements, ordered according to the
HS 2-9-1-8
decreasing contents of the alloying
W=2%
Mo=9%
V=1%
Co=8%
br-er05-09.cdr
elements and numbers, which in the
© ISF 2004
Codes according to the chemical composition
sequence of the designating alloying elements give reference about their content. The individual numbers stand
Figure 5.9
for the medium content of the respective alloying element, the content had been multiplied by the factor as indicated in Fig. 5.9/Table 5.1 and rounded up to the next whole number.
5. Welding Plain and Low Alloy Steels
53
The alloyed steels are labelled with the code letter X, a number which again complies with the hundredfold of the mean value of the range stipulated for the carbon content, the chemical symbols of the alloying elements, ordered according to decreasing contents of the elements and numbers which in sequence of the designating alloying elements refer to their content. High-speed steels are designated with the code letter HS and numbers which, in the following sequence, indicate the contents of elements:: tungsten (W), molybdenum (Mo), vanadium (V) and cobalt (Co).
The European Standard DIN EN 10027-2 (September 1992) specifies a numbering system for the designation of steel grades, which is also called material number system.. The structure of the material number is as follows: 1.
XX
XX (XX) Sequential number The digits inside the brackets are intended for possible future demands. Steel group number (see Fig. 5.10) Material main group number (1=steel)
5. Welding Plain and Low Alloy Steels
Figure 5.10 specifies the material numbers for the material main group „steel“.
Figure 5.10
54
5. Welding Plain and Low Alloy Steels
55
The influence of the austenite grain size on the transformation behaviour has been explained in Chapter 2. Figure 5.11 shows the dependence between grain size of the austenite which develops during the welding cycle, the distance from the fusion line and the energy-per-unit length from the welding method. The higher the energy-peruntil
length,
the
bigger the austenite grains in the
13
HAZ and the width
Austenite grain size index according to DIN 50601
Energy-per-unit length in kJ/cm
11 9
12
18
of
36
the
creases.
9
HAZ
in-
Such
coarsened austen-
7
ite grain decreases 5
the critical cooling 3 0
0,2
0,4 0,6 Distance of the fusion line
0,8
mm
1,0 © ISF 2004
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Influence of the energy-per-unit length on the austenite grain size
time, thus increasing the tendency of the steel to harden.
Figure 5.11
With fine-grained structural steels it is tried to suppress the grain growth with alloying elements. Favourable are nitride and carbide forming alloys. They develop precipitations which suppress undesired grain growth. There is, however, a limitation due to the solubility of these precipitations, starting with a certain temperature, as shown in Figure 5.12. Steel 1 does not contain any precipitations and shows therefore a continuous grain growth related to temperature. Steel 2 contains AIN precipitations which are stable up to a temperature of approx. 1100°C, thus preventing a growth of the austenite grain.
5. Welding Plain and Low Alloy Steels
56
With
Grain size index according to DIN 50601
mm 1 8 6
Medium fibre length
4
2
10 8 6
-1
4
2
-2
temperatures,
precipitations dissolve and cannot
-2
suppress a grain growth any more.
0
Steel 3 contains mainly titanium car-
2
bonitrides of a much lower grain-
4
refining effect than that of AIN. Steel 4
6
is a combination of the most effective properties of steels nos. 2 and 3.
8
Steel 1 Steel 2 Steel 3 Steel 4
6 10-3 12 900
1000
1100 1200 Austenitization temperature
1300
The importance of grain refinement for the mechanical properties of a
°C
1400
steel is shown in Figure 5.13. Pro-
Steel
%C
% Mn
% Al
%N
% Ti
1
0,21
1,16
0,004
0,010
/
2
0,17
1,35
0,047
0,017
/
3
0,18
1,43
0,004
0,024
0,067
4
0,19
1,34
0,060
0,018
0,140
vided the temperature keeps constant, the yield strength of a steel increases with decreasing grain size.
© ISF 2004
br-er05-12.cdr
This influence on the yield point Rel is
Austenite grain size as a function of the austenitization temperature
specified
According
to
the
1 d
900
the yield point is propor-
tional to the root of the medium grain
N/mm² 800
Yield point or 0,2 boundary
law, the increase of
Temperature in °C:
700
-193 -185
600
-170 -155
-100
σi
300
-40
stands for the inter-
200
diameter d.
-180
500 400
+20 0
nal friction stress of
1
2
3
4
5 6 -1/2 Grain size d
7
grain
for
is
a
mm
-1/2
10
Connection between yield point and grain size
boundary
resistance K measure
The
8
© ISF 2004
br-er-05-13.cdr
material.
Hall-Petch-law:
the
above-mentioned
inversely
in
Rel = σ i + K ⋅
Figure 5.12
the
these
-4
10
10 8
higher
Figure 5.13
the
influence of the grain size on the forming mechanisms. Apart from this increase of the yield point, grain refinement also results in improved toughness values. As far as
5. Welding Plain and Low Alloy Steels
57
structural steels are concerned, this means the improvement of the mechanical properties without any further alloying. Modern fine-grained structural steels show improved mechanical properties with, at the same time, decreased content of alloying elements. As a consequence of this chemical composition the carbon equivalent decreases, the weldability is improved and processing of the steel is easier. The major advanSteel type Stahlsorte
S235JR (St37-2)
S355J2G3 (St52-3)
S690Q (StE690)
S890Q (StE890)
S960Q (StE960)
Ratio Verhältnis S235JR - S960Q
N/mm2
215
345
690
890
960
1:5
Plate thickness Blechdicke
mm
50
31
14,4
11
10
5:1
Yield point Streckgrenze Weld cross-section Nahtquerschnitt
mm2
870
370
100
60
50
17 : 1
Welding wire Øø1.2 Schweißdraht 1.2
mm
SG2
SG3
NiMoCr
X 90
X 96
-
Welding wire costs Schweißdrahtkosten
Ratio Verhältnis
1
1
2,4
3,2
3,3
1 : 3,3
Steel costs Stahlkosten
Ratio Verhältnis
1
1,2
1,9
2,3
2,4
1 : 2,4
Weld metal costs Schweißgutkosten
Ratio Verhältnis
5,3
2,3
1,5
1,16
1
5,3 : 1
Special weld costs Spez. Schweißnahtkosten
Ratio Verhältnis
12
5,1
1,8
1,18
1
12 : 1
Costs ratio inclusive base Kostenverhältnis inklusive materials Grundwerkstoffe
Boundary condition: Randbedingungen:
tages of microalloyed
fine-grained
structural steels in comparison
with
conventional structural
5:1
steels
shown
welding process = MAG Schweißverfahren = MAG
in
are
Figure
Deposition rate = 3 kg=welding wire/h, weld /shape X -60° X - 60° Abschmelzleistung 3 kg Schweißdraht h, Nahtform
5.14. Due to the
Costs labour and equipment == 60 30€/h Lohn-ofund Maschinenkosten DM / h Special costs = weld filler materials + welding Spez. weld Schweißnahtkosten = Schweißzusatzwerkstoffe + Schweißen
considerably better
Berechnungsgrundlage =szul = Re / 1.5 Calculation base = szul = Re/1.5 © ISF 2004
br-er-05-14.cdr
mechanical proper-
Influence of the steel selection on the producing costs of welded structures
ties of the finegrained
Figure 5.14
structural
steel in comparison with unalloyed structural steel, substantial savings of material are possible. This leads also to reduced joint cross-sections and, in total, to lower costs when making welded steel constructions. Based
on
steels
the
alloyed
unalloyed
classification Figure
5.2,
of Fig-
low-alloyed mild steel
higher-carbon steel Hardening Underbead cracking
ure 5.15 divides the steels with regard
rimmed steel
to their problematic
cutting of segregation zones
processes
during
welding. When it
killed steel duplex killed steel
cold brittleness (coarse-grained recrystallization after critical treatment) stress corrosion cracking safety from brittle fracture
comes to unalloyed
high-alloyed
hardening corrosion tool steels special properties are resistant steels achieved, for example: Hardening, special properties heat resistance, are achieved tempering resistant, high-pressure hydrogen resistance, toughness at low temperatures, surface treeatment condition, etc. ferritic
pearlitic-martensitic
austenitic
grain increase in the weld interfaces
hardening embrittlement formation of chromium carbide
grain desintegration stress corrosion cracking hot cracks (sigma phase embrittlement)
Post-weld treatment for highest corrosion resistance © ISF 2004
br-er-05-15.cdr
steels, only ingot
Classification of steels with respect to problems during welding
Figure 5.15
5. Welding Plain and Low Alloy Steels
58
casts, rimmed and semi-killed steels are causing problems. “Killing” means the removal of oxygen from the steel bath. Figure 5.16 shows cross-sections of ingot blocks with different oxygen contents. Rimming steels with increased oxygen content show, from the outside to the inside, three different zones after solidification: 1.: a pronounced, very pure outer envelope, 2.: a typical blowhole formation (not critical, blowholes are forged together during rolling), 3.: in the centre
a
segregated
clearly zone
where unfavourable elements like sulphur and phosphorus are enriched.
0,025 0,012
During rolling, such
0,003
fully killed steel
semi-killed steel
zones are stretched
rimmed steel
along the complete
Figures: mass content of oxygen in % © ISF 2004
br-er-05-16.cdr
length of the rolling
Ingot cross-sections after different casting methods
profile. Figure 5.16
Figure 5.17 shows important points to be observed during welding such steels. Due to their enrichment with alloy elements, the segregation zones are more transformation-inert than the outer
envelope
a
b
and are inclined to hardening.
In
addition, they are sensitive
to
cracking,
as,
hotin
B
these zones, the
D
C
E
elements phosphorus
and
sulphur
© ISF 2004
br-er-05-17.cdr
are
enriched.
Example of unfavourable (a) and favourable (b) welds
Figure 5.17
5. Welding Plain and Low Alloy Steels
59
Therefore, “ touching” such segregation zones during welding must be avoided by all means. In the case of lowalloy
steels,
the
Microstructures
Average Brinell Hardness (Approximately)
Ferrite
80
Austenite
250
Perlite (granular)
200
welding
Perlite (lamellar)
300
observed.
Sorbite
350
Troostite
400
Cementite
600 - 650
hardness values of
Martensite
400 - 900
various microstruc-
problem
of
hardening
during must
be Fig-
ure 5.18
shows
tures. The highest
© ISF 2004
Br-er-05-18.cdr
HAZ
hardness
Hardness of Several Microstructures
values
can be found with Figure 5.18
martensite
and
cementite. Hardness values of cementite are of minor importance for unalloyed and low-alloy steels because its proportion in these steels remains low due to the low Ccontent. However, hardening because of martensite formation is of greatest importance as the martensite proportion in the microstructure depends mainly on the cooling time. Figure 5.19 shows the essential influHV
HRC
root cracking presumable
400
41
1290
70
root cracking possible
400 - 350
41 - 36
1290 - 1125
70 - 60
no root cracking
350
36
1125
60
sufficient operational safety without heat treatment
280
28
900
30
ence of the martensite
content
in
the HAZ on the crack formation of welded
joints.
Hardening through martensite
forma-
with maximum martensite content %
strength, calculated at max. hardness N/mm2
maximum hardness
If too much martensite develops in the heat affected zone during welding (below or next to the weld), a very hard zone will be formed which shows often cracks.
tion is not to be © ISF 2004
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expected with pure
Influence of Martensite Content
carbon steels up to about
0,22%,
Figure 5.19
5. Welding Plain and Low Alloy Steels
60
because the critical cooling rate with these low C-contents is so high that it normally won’t be reached within the welding cycle. In general, such steels can be welded without special problems (e.g., S. 235). In addition to carIIW
C - Äqu. = C +
Mn Cr + Mo + V Cu + Ni + + 6 5 15
Stout
C - Äqu. = C +
Mo Ni Cu Mn Cr + Mn + + + 6 10 20 40
Ito and Bessyo
PCM = C +
Mannesmann
C - Äqu.PLS = C +
Hoesch
C - Äqu. = C +
C ET
Thyssen
bon, all other alloy elements are important
Si Mn + Cu + Cr Ni Mo V + + + + + 5B 30 20 60 15 10
site
formation
in
the welding cycle,
Si + Mn + Cu + Cr + Ni + Mo + V 20
as they have sub-
Mn + Mo Cr + Cu Ni = C+ + + 10 20 40
stantial
PLS = pipeline steels
it
comes to marten-
Si Mn + Cu Cr Ni Mo V + + + + + 25 16 20 60 40 15
C-Äqu.= carbon equivalent (%)
when
influence
on the transforma-
PCM = cracking parameters (%) © ISF 2002
Br-er-05-20.cdr
tion behaviour of Definition of C - Equivalent
steels
(see
Fig. 2.12 ). It is not
Figure 5.20
appropriate just to take the carbon content as a measure for the hardening tendency of such steels. To estimate the weldability, several authors developed formulas for calculating the so-called carbon equivalent, which include the contribution of the other alloy elements to hardening tendency, (Fig. 5.20). As these approximation formulas are empirically determined as
for
0,35
Tp ==750 CET - 150- 150 Tp 750 CET
delta Tp HD HD0,35 - 100 delta Tp= 62 = 62 - 100 80
200
the
delta Tp [°C]
and
100
250
hardening tendency
Tp [° C]
150
100
d = 30 mm d = 30 mm HD HD = 4= 4 1 kJ/mm Q = Q1=kJ/mm
0 0,2
tions
like
0,3
0,4
CET = =0,33 % CET 0,33 % = 30mm mm d =d30 kJ/mm Q =Q1= 1kJ/mm
0 0
0,5
5
60
heat
10
15
20
25
Wasserstoffgehalt Hydrogen contentHD of des theSchweißgutes weld metal [%]
Kohlenstoffäquivalent CET [%] Carbon aquivalent
plate
40
delta TpTp = 160 tanhtanh (d/35)(d/35) - 110 - 110 delta = 160
thickness,
40
20
50
the general condi-
60
delta Tp CETCET - 32)-Q32) - 53Q CET + 32 delta Tp= (53 = (53 - 53 CET + 32 20
50
CET = 0,2 %
CET = 0,2 %
CET = 0,4 %
CET = 0,2 %
CET = 0,4 %
CET = 0,2 %
0
delta Tp [°C]
input, etc., are also
delta Tp [°C]
40
30
-20
-40
20 -60
of importance, the
10
CET 0,4 CET ==0,4 %% HD = 2 2 HD QQ== 11kJ/mm kJ/mm
0
carbon
equivalent
cannot be a com-
0
20
40
60
80
100
-80
d =d50 = 50mm mm HDHD = =8 8
-100 0
0,5
Tp =697 CET + 160 tanh (d/35) + 62 HD br-er05-21.cdr
mon limit value for the weldability. For the determina- Figure 5.21
1
1,5
2
2,5
3
3,5
4
4,5
Wärmeeinbringen Heat input Q [kJ/mm]
Plate thickness Blechdicke d [mm]
0,35
+ (53 CET - 32) Q - 328
Source: Quelle: DIN EN 1011-2
Calculation of the preheating temperatures
© ISF 2005
5
5. Welding Plain and Low Alloy Steels
61
tion of the preheating temperature Tp, the formula as shown in Fig. 5.21 is used. The effects of the chemical composition which is marked by the carbon equivalent CET, the plate thickness d, the hydrogen content of the weld metal HD and the heat input Q are considered. The essential factor to martensite forma-
Temperature T
Tmax
tion in the welding cycle is the cooling
°C
time. As a measure 800
of cooling time, the DT
time of cooling from 500
800 to 500°C (t8/5) is
t8/5
defined (Fig. 5.22). t800
t500
s
The
Time t
temperature
© ISF 2004
br-er-05-22.cdr
range was selected
Definition of t8/5
in such a way that it covered the most
Figure 5.22
important structural transformations and that the time can be easily transferred to the TTT diagrams. Figure 5.23
shows 2000
measured
time-
temperature
distri-
°C
ity of a weld. Peak values
and
dwell
times depend obvi-
Temperature T
butions in the vicin-
B
1500
A
A
of
the
B
500 C
0 0
measurement
10mm
1000
ously on the location
and
50
100
150
200
250
s
300
Time t © ISF 2004
br-er-05-23.cdr
are clearly strongly determined by the heat
C
conduction Figure 5.23 conditions.
Temperature-time curves in the adjacence of a weld
5. Welding Plain and Low Alloy Steels
62
With the use of thinner plates with complete heating of the cross-section during welding, the heat conductivity is only carried out in parallel to the plate surface, this is the two-dimensional heat dissipation. With thicker plates, e.g. during welding of a blind bead, heat dissipation can also be carried out in direction of plate thickness, heat dissipation is three-dimensional.
3 - dimensional:
These two cases
K3 t8 / 5 =
universal formula:
extended formula For low-alloyed steel:
ö h U ×I æ 1 1 ÷ × ×ç 2 × p × l v çè 500 - T0 800 - T0 ÷ø
) Uv× I × æçç 5001- T
(
t8 / 5 = 0,67 - 5 ×10 - 4 T0 ×
è
-
0
are covered by the formulas given in
ö 1 ÷ ×h ¢ × N 3 800 - T0 ø÷
Figure 5.24, which K2
2 - dimensional: t8 / 5 =
universal formula:
extended formula For low-alloyed steel:
provide a method
2 2 2 ö ù ö æ h2 1 1 æ U × I ö 1 éæç ÷ ú ÷ -ç ×ç ÷ × ×ê 4 × p × l × r × c è v ø d 2 êçè 500 - T0 ÷ø çè 800 - T0 ÷ø ú û ë
of calculating the
2 2 2 ö æ ö ù 2 1 1 æ U × I ö 1 éæç ÷ -ç ÷ ú ×h ¢ × N 2 t8 / 5 = 0,043 - 4,3 ×10 -5 T0 × ç ÷ × 2 ×ê è v ø d ëêçè 500 - T0 ÷ø çè 800 - T0 ÷ø ûú
(
formula for the transition thickness of low-alloyed steel:
)
dü =
0,043 - 4,3 ×10 -5 T0 U ×I ×h ¢ × 0,67 - 5 ×10 - 4 T0 v
cooling time t8/5 of low-alloyed steels.
æ ö 1 1 ÷÷ × çç + è 500 - T0 800 - T0 ø
In the case of a © ISF 2004
br-er-05-24.cdr
three-dimensional
Calculation equation for two- and three-dimensional heat dissipation
heat
dissipation,
t8/5 it independent
Figure 5.24
of plate thickness. In the case of two-dimensional heat dissipation it is clear that t8/5 becomes the shorter the thicker the plate thickness d is. Provided, the cooling times are equal, the plate thickness can be calculated from these relations where a two-dimensional heat dissipation changes to a three-dimensional heat dissipation. Figure 5.25 shows welding methods
the influence of the
TIG-(He)-welding
welding method on
TIG-(Ar)-welding
the heat dissipa-
MIG-(Ar)-welding
tion. With the same
MAG-(CO2)- welding
heat
the
Manual arc welding
is
SA welding
input,
energy
which
0
transferred to the
0,1
0,2
0,3
0,4
0,5
0,6
0,7
0,8
0,9
Relative thermal efficiency degree h‘
base
material
© ISF 2004
Br-er-05-25.cdr
depends
on
the
Relative thermal efficiency degree of different welding methods
Figure 5.25
1
5. Welding Plain and Low Alloy Steels
63
welding method. This dependence is described by the relative thermal efficiency ŋ’. The influence of the
groove
Type of weld
ge-
2-dimensional heat dissipation
ometry is covered
weld factor 3-dimensional heat dissipation
1
1
0,45 - 0,67
0,67
0,9
0,67
0,9
0,9
by seam factors according
to
Fig. 5.26. Empirically determined, these factors were introduced for an
© ISF 2004
br-er-05-26.cdr
easier calculation.
Weld factors for different weld geometries
For other groove geometries, tests to
measure
Figure 5.26
the
cooling time are recommended. Fig. 5.27 shows the transition of the two-dimensional to the three-dimensional heat dissipation for two different preheating temperatures in form of a curve according to the equation of Fig. 5.24. Above the curve, t8/5 depends only on the energy input, but not on the plate thickness, heat dissipation is carried out three-dimensionally.
5
cooling time t8/5 [s] 10 15 20
cm
cooling time t8/5 [s] 10 20 30
25
Plate thickness
TA=20°C
40
50
TA=200°C
60 80 100 150
3
30 40
3-dimensional 2
3-dimensional
60 100
1
2-dimensional
2-dimensional 0 0
10
20
30
40
50
0
10
20
30
40
Heat input E.h.Nn [kJ/cm] © ISF 2004
Br-er-05-27.cdr
Transition From Two to Three Dimensional Heat Flow
Figure 5.27
50
5. Welding Plain and Low Alloy Steels
64 Fig. 5.28 shows the possible range of heat input depending on the electrode diameter. It is clear that a relatively large working range is available for
arc
welding
procedures. variation
of
A the
energy-per-unit Figure 5.28
length
can
be
carried out by alteration of the welding current, the welding voltage and the welding speed.
Fig. 5.29 depicts variations of the heat input during manual metal arc welding. The shorter the fused electrode distance, i.e., the shorter the extracted length, the higher the energyper-unit length.
Figure 5.29
5. Welding Plain and Low Alloy Steels
65
In order to minimize calculation efforts in practice, the specified relations were transferred into nomograms from which permissible welding parameters can be read out, provided some additional data are available. Fig. 5.30 shows diagrams for twodimensional heat dissipation, where a dependence between energy-per-unit length, cooling time and preheating temperature is given, depending on the plate thickness. .
50 40 30
T0 200°C 150°C 100°C
20
20°C
Cooling time t8/5 in s
10
d = 7,5 mm
7 50 40 30
T0 200°C 150°C 100°C
20
20°C
10
d = 10 mm
7 50 40 30
T0 200°C 150°C 100°C
20
20°C
10
d = 15 mm
7 50 40 30
T0 200°C 150°C 100°C
20 transition to 3-dimensional heat flow
10
20°C d = 20 mm
7 5
6
br-er05-30.cdr
7 8 9 10
15 20
30
kJ/cm 50
Heat input E
© ISF 2004
Dependence of E, t8/5 and d During SA - Welding
Figure 5.30
If a fine-grained structural steel is to be welded, the steel manufacturer presets a certain interval of cooling times, where the steel characteristics are not too negatively affected. The user lays down the plate thickness and, through the selection of a welding method, a specified range of heat input E. Based on the data E and t8/5 the diagram provides the required preheating temperature for welding the respective plate thickness.
5. Welding Plain and Low Alloy Steels
With the transition to thicker plates,
Transition thickness dÜ
50 mm 40
the diagrams in Fig. 5.31 apply. The
aera of 3-dimensional heat flow
30
T0
20 15
10 9 8 7
66
0 °C °C 20 °C 250 00 1 °C 150 C ° 20
upper part of the figure determines whether a two-dimensional or a threedimensional heat dissipation is pre-
area of 2-dimensional heat flow
sent. For the three-dimensional heat dissipation, the lower diagram applies
5
6
7 8 9 10
15 20
30
kJ/cm 50
where the same information can be
Heat input E 50 s 40
determined,
Cooling time t8/5
of
thickness, as with Fig. 5.30.
30
20 15
0
25 T
0
°C 0
20
°C 0
15
°C 0
10
10 9 8 7
independent
5
6
7 8 9 10
°C 20
°C
15 20
30
Heat input E
br-er05-31.cdr
kJ/cm 50 © ISF 2004
Dependence of E, T0, t8/5 And dÜ
Figure 5.31
The
relation
be-
tween current and
35 V
voltage for MAG
gas composition: C1 100% CO2 M21 82% Ar + 18% CO2 M23 92% Ar + 8% O2
C1 M21
30
in Fig. 5.32
and
the used shielding gas is one of the
Welding voltage
M23
welding is shown
25
20
15
parameters. Welding
voltage
contact tube distance ~15mm
150
and
welding current, or
3,5 br-er-05-32.cdr
wire feed speed,
contact tube distance ~19mm
200
250 Welding current
A
300
5,5
7,0 Wire feed
9,0
10,5
8,0
m/min © ISF 2004
Dependence of Current And Voltage During MAG-Welding, Solid Wire, Æ1.2 mm
determine the type of arc.
4,5
spray arc
mixed arc
short arc
Figure 5.32
plate
5. Welding Plain and Low Alloy Steels
67
The diagram in Fig. 5.33 demonF3 = 0,67 F2 = 0,67
h'UP = 1 h'MAG = 0,85 dU max = 32 mm dU min = 15 mm
t8/5 max = 30 s t8/5 min = 6 s
Emax = 66 kJ/cm Emin = 14 kJ/cm
ness, heat input E and cooling time
60 fillet welds T0= 150 °C
kJ/cm
30s
50
t8/5
kJ/cm
temperature of T0 = 150°C. If d and t8/5 are given, the acceptable range of
25s
toughness affection
53 47 20s
35
41
30
15s
25
heat input can be determined with the
Heat input E MAG - weldind
40
35
help of this diagram. The kinks of the curves mark the transition between
29 10s
20 15
for fillet welds at a preheating
70
59
45
Heat input E SA - welding
strates the dependence of plate thick-
two-dimensional
and
three-
23
dimensional heat dissipation.
18 6s
10
12 cracking tendency
5 0
0
5
10
15
20 25 30 Plate thickness
6 mm
0 40
br-er05-33.cdr
© ISF 2004
Permissible E-Range During SA - And MAG - Welding
Figure 5.33 Fig. 5.34 shows the same dependt8/5 max = 30 s t8/5 min = 6 s
Emax = 49 kJ/cm Emin = 10 kJ/cm
70
60 butt welds T0= 150 °C
kJ/cm
kJ/cm 59
50 toughness affection
53
45 30s
47
40 25s
35 30
20s
41 35 29
25 15s
20
23
15
10s
18
10
6s
12
cracking tendency
5 0
Heat input E MAG - welding
preparation..
F3 = 0,9 F2 = 0,9
h'UP = 1 h'MAG = 0,85 dU max = 34 mm dU min = 15 mm
Heat input E SA - welding
ence for butt welds with V groove
0
br-er05-34.cdr
5
10
15
20 25 30 Plate thickness
mm
6 0 40 © ISF 2004
Permissible E-Range During SA - And MAG - Welding
Figure 5.34
5. Welding Plain and Low Alloy Steels
68
The curve family in Fig. 5.35 shows the dependence of the heat input from the welding speed as well as the acceptable working range. The parameters of the curves 1 to 8 in the table
25 kJ/cm
2
3
4
5
6
7
8
V
29
27
24
22
20
19
18
17
A
300 275 250 225 200 175 150 125
vZ(m/min) 10.5 9.0 8.0 7.0
20
from Figures 5.32 and 5.34 and apply
5.5 4.5 3.5 3.0
1
2
Heat input E
have been taken
1
curve
only
wor king ra
3
15
nge
4
5
related
conditions like wire
6
7
10
for
8
diameter,
5
wire
feed,
0 10
15
20
25
30 35 40 45 Welding speed vS
50 cm/min 60
welding
voltage, etc.
MAG/ M21 (82% Ar, 18% CO) br-er-05-35.cdr
© ISF 2004
E as a Function of Welding Speed, Solid Wire, Æ1.2mm
Figure 5.35
shows
Sheet
Nr. 0916). In this example, a plate thickness of 15 mm and a cooling
time
t8/5
be-
1
2
3
4
5
6
7
8
V
29
27
24
22
20
19
18
17
59
A
300 275 250 225 200 175 150 125
toughness affection
53
45 30s
47
40 25s
35 30
20s
15s
20
10s
15 10
6s
cracking tendency
5 0
41 35 29
25
0
5
10
15
20 25 30 Plate thickness
mm
curve
kJ/cm
23 18
16 12 13 6 0 40
vZ(m/min) 10.5 9.0 8.0 7.0
5.5 4.5 3.5 3.0
25 kJ/cm 20
1
2
heat input E
Reference
SA - welding
(according to DVS-
butt welds T0= 150 °C
50
Heat input E
for such diagrams
70
60 kJ/cm
MAG - welding
a reading example
Heat input E
Figure 5.36
16 15 13
work
ing
3
4
rang
e
5
6
7
10
8
5
33
0 10
15
20
25
41
30 35 40 45 Welding speed vS
50 cm/min 60
© ISF 2004
br-er-05-36.cdr
Determination of Welding Speed for MAG - Welding
tween 10 and 20 s are given. In this case, the maximum
Figure 5.36
cooling time for MAG welding is 15 s. A solid wire with a diameter of 1.2 mm at 29V and 300A is used. The left diagram provides heat input values between 13 and 16 kJ/cm, based on the given data. Using these values, the acceptable range of welding speeds can be taken from the diagram on the right.
5. Welding Plain and Low Alloy Steels
69
Fig. 5.37 presents a simplification of
800 °C
the determination of the microstruc-
700
tural composition and cooling time subject to peak temperatures which
Temperature
F
occur in the welding cycle. In the
line. The point of intersection of the
500 400
M
Peak temperature 1000°C 1400°C
200
HV30=400
200
300
1400
Peak temperature
the point of heat input at the lower
P B
300
lower diagram, the point of the plate thickness at the top line is linked with
600
°C
1000 Arc3
800
Arc1
linking line with the middle scale 600
represents the cooling time t8/5 .
middle diagram in which transition field the final microstructures are
1
plate thickness 40
If the peak temperature of the welding cycle is known, one can read from the
F+P
F+B
B+M
M
1200
30
25
three-dimensional
two-dimensional
10
1
20
s
100
15
10 9 8
7
6
1000
t8/5
5 mm 4
300 200 100
2 3
5
10
20
50 100 200 400 s 1000 0
100 °C
200
t8/5
preheating temperature
energy-per-unit length 6
8
10
20
30
40
50 kJ/cm 70
bie5-37.cdr
formed. The advantage of the determination of microstructures compared
© ISF 2004
Peak temperature/cooling time – diagram for the determination of t8/5 and the structure
with the upper TTT diagram is that Figure 5.37 a TTT diagram applies only for exactly one peak temperature, other peak temperatures are disregarded. The disadvantage of the PTCT diagram (peak temperature cooling time diagram) is the very expensive determination, therefore, due to the measurement efforts a systematic application of this concept to all common steel types is subject to failure.
6. Welding High Alloy Steels
6. Welding High Alloy Steels
71
Basically stainless steels are characterised by a chromium content of at least 12%. Figure 6.1 shows a classification of
corrosion
corrosion-resistant steels
resistant
steels. They can be sin-
scale- and heat-resistant steels
stainless steels
gled out as heat- and scale-resistant
and
stainless steels, depend-
perlitic martensitic
semi-ferritic
ferritic
ferritic-austenitic
X40Cr13
X10Cr13
X8Cr13
X20CrNiSi25-4
austenitic
ing on service temperature. Stainless steels are used at room temperature conditions and for water-
non-stabilized
stabilized
(austenite with delta-ferrite) X12CrNi18-8
(austenite without delta-ferrite) X8CrNiNb16-13 © ISF 2002
br-er-06-01e.cdr
based media, whilst heatClassification of Corrosion-Resistant Steels
and scale-resistant steels are applied in elevated
Figure 6.1
temperatures and gaseous media. Depending on their microstructure, the alloys can be divided into perlitic-martensitic, ferritic, and austenitic steels. Perlitic-martensitic steels have a high strength and a high wear resistance, they are used e.g. as knife steels. Ferritic and corrosion resistant steels are mainly used as plates for household appliances and other decorative purposes. The most important group are austenitic steels, which can be used for very many applications and which are corrosion resistant against most media. They have a very high low temperature impact resistance. Based on the simple Fe-C T
T A4
T d
phase diagram (left figure), d
Figure 6.2 shows the ef-
A4
A4 g
g
A3
g
a(d)
fects
of
two
different
A3
A3
groups of alloying elements
a a
on the equilibrium diagram. Alloy elements in %
Alloy elements in % Chromium Vanadium Molybdenum Aluminium Silicon
Alloy elements in %
Ferrite
Nickel Manganese Cobalt
developers
with
chromium as the most important element cause a © ISF 2002
br-er-06-02e.cdr
Modifications to the Fe-C Diagram by Alloy Elements
Figure 6.2
strong reduction of the aus-
6. Welding High Alloy Steels
72
tenite area, partly with downward equilibrium line according to Figure 6.2 (central figure). With a certain content of the related element, there is a transformation-free, purely ferritic steel. An opposite effect provide austenite developers. In addition to carbon, the most typical member of this group is nickel.
Carbon l l l Chromium l Nickel l l
Steel type, no. All types l l l All types l
Effect Increases the strength, supports development of precipitants which reduce corrosion resistance, increasing C content reduces critical cooling rate Works as ferrite developer, increases oxidation- and corrosion-resistance
All types
Works as austenite developer, increases toughness at low temperature, grain-refining
Works as strong austenite developer (20 to 30 times stronger than Nickel) 1.4511,1.4550, Binds carbon and decreases tendency to Niobium 1.4580 u.a. intergranular corrosion l All types Increases austenite stabilization, reduces hot Manganese l l crack tendency by formation of manganese l l sulphide Improves creep- and corrosion-resistance Molybdenum 1.4401,1.4404, 1.4435 and others. against reducing media, acts as ferrite l developer l l 1.4005, 1.4104, Phosphorus, Improve machinability, lower weldability, 1.4305 selenium, or l reduce slightly corrosion resistance l sulphur l All types Improves scale resistance, acts as ferrite l Silicon developer, all types are alloyed with small l l contents for desoxidation l 1.4510, 1.4541, Binds carbon, decreases tendency to Titanium l 1.4571 and others intergranular corrosion, acts as a grain refiner l l and as ferrite developer Type 17-7 PH Works as strong ferrite developer, mainly Aluminium l l used as heat ageing additive Type 17-7 PH, Copper Improves corrosion resistance against certain l 1.4505, 1.4506 media, decreases tendency to stress l l corrosion cracking, improves ageing l l Oxygen l
Austenite developers cause an extension of the austenite area to Figure 6.2 (left figure) and form a purely austenitic and transformation-free steel.
Special types l
The table in Figure 6.3 summarises the effects of some selected elements on high alloy steels.
1800 °C
Effects of Some Elements in Cr-Ni Steel
Figure 6.3
S+a
1400
© ISF 2002
br-er06-03e.cdr
S
1600
Temperature
Element
1200 g
a
g+a
1000 800 d+a
d+a'
d
600 a'
a
The binary system Fe-Cr in Figure 6.4 shows
400 200
the influence of chromium on the iron lattice. Starting with about 12% Cr, there is no more
0 Fe
10
20
30
40
50
60
70
80
90 % Cr
Chromium
transformation into the cubic face-centred lattice, the steel solidifies purely as ferritic. In
© ISF 2002
br-er06-04e.cdr
Binary System Fe - Cr
the temperature range between 800 and 500°C this system contains the intermetallic σ-phase, which decomposes in the lower
Figure 6.4
temperature range into a low-chromium α-solid solution and a chromium-rich α’-solid solution. Both, the development of the σ-phase and of the unary α-α’-decomposition cause a strong
6. Welding High Alloy Steels
73
embrittlement. With higher alloy steels, the diffusion speed is greatly reduced, therefore both processes require a relatively long dwell time. In case of technical cooling, such embrittlement processes are suppressed by an increased cooling speed. Nickel is a strong austenite developer, see Figure 6.5 Nickel and iron develop in this system under elevated temperature a complete series of face-centred cubic solid solutions. Also in 1600 °C 1400
d
Fe Ni3
the binary system Fe-Ni S+d
S+g
decomposition
d+g
in the lower temperature
1200
range take place.
g
Temperature
processes
1000
Along two cuts through the
800
ternary system Fe-Cr-Ni, 600
a
Figure 6.6 shows the most
a+g
400
Fe Ni3
important
200
phases
which
develop in high alloy steels.
0 Fe
20
10
30
50
40
60
70
80
90 % Ni
Nickel
br-er-06-05e.cdr
© ISF 2002
A solidifying alloy with 20%
Binary System Fe - Ni
Cr and 10% Ni (left figure) forms at first δ-ferrite. δ-
Figure 6.5
ferrite is, analogous to the 70 % Fe
60 % Fe
1600 °C
1600 °C
S
1500
S+g
S+d
1400
S+d+g
1400
from the melt solidifying
S+g
S+d
body-centred
1300
1200 g
d+g
cubic
solid
solution. However α-ferrite
1200 d
g
d+g
d
1100
1100
is developed by transfor-
1000
1000
mation of the austenite, but
900
900
800
800
700
d+s
d+g+s
is of the same structure g+s
d+s
g+s
from the crystallographic
d+
g+
s
Temperature
1300
Fe-C diagram, the primary
S
1500
S+d+g
700
0
5
10
15
30
25
20
15
20 % Ni 10 % Cr
0
5
10
15
20
40
35
30
25
20
% Ni
15
% Cr
point of view, see Figure
© ISF 2002
br-er-06-06e.cdr
Sections of the Ternary System Fe-Cr-Ni
Figure 6.6
25
6.4.
6. Welding High Alloy Steels
74
During an ongoing cooling, the binary area ferrite + austenite passes through and a transformation into austenite takes place. If the coolls
ing is close to the equilibrium, a partial transst ee iti c-
st ee
takes place in the temperature range below
Au s
te n
c te ni ti
800°C. Primary ferritic solidifying alloys show
3.
4.
Au s
ar M 2.
fe rri tic
ls
s st ee l
s te n
si tic
st ee l Fe rri tic 1.
formation of austenite into the brittle α-phase
C
£ 0.1
0.1 1.2
£ 0.1
£ 0.1
Si
max. 1.0
max. 1.0
max. 1.0
max. 1.0
Mn
max. 1.0
max. 1.5
max. 2.0
max. 2.0
Cr
15 18
12 18
17 26
24 28
Mo
up to 2.0
up to 1.2
up to 5.0
up to 2.0
Ni
£ 1.0
£ 2.5
7 26
4 7.5
a reduced tendency to hot cracking, because δ-ferrite can absorb hot-crack promoting elements like S and P. However primary austenitic solidifying alloys show, starting at a certain alloy content, no transformations during cool-
up to 2.2
Cu Nb
+
+
Ti
+
+
Al
+
+
ing (14% Ni, 16% Cr, left figure). Primary austenitic solidifying alloys are much more susceptible to hot cracking than primary fer-
+
V N
+ indicates that the alloy elements can be added in a defined content to achieve various characteristics
+ +
S
ritic solidifying alloys, a transformation into the
+
σ-phase normally does not take place with
© ISF 2002
br-er06-07e.cdr
Typical Alloy Content of High-Alloy Steels
these alloys. Figure 6.7 shows some typical compositions
Figure 6.7
of certain groups of high alloy steels.
The diagram of Strauß and Maurer in Figure 6.8 shows the influence on the microstructure formation of steels with a C-content of 0,2%. The classification of high-alloy steels in Figure 6.1 is based on this dia-
28
gram. If a steel only con-
% 24
tains C, Cr and Ni, the austenite
Nickel
20
lowest austenite corner will
16
be at 18% Cr and 6% Ni.
12
And also other elements
8
austen it
4
ensite
martensite / troostite / sorbite ferrite / perlite
0
e / mart
0
2
4
austenite / ferrite austenite
/ martens
ite / ferrite
martensite / ferrite 6
8
10
12 14 Chromium
16
18
20
22 © ISF 2002
br-er-06-08e.cdr
Maurer - Diagram
24 % 26
than Ni and Cr work as an austenite or ferrite developer.
The
these
elements
is
of de-
scribed by the so-called chromium
Figure 6.8
influence
and
nickel
6. Welding High Alloy Steels
75
equivalents. The Schaeffler diagram reflects additional alloy elements, Figure 6.9. It represents molten weld metal of high alloy steels and determines the developed microstructures after cooling down from very high temperatures. The diagram was always prepared considering identical cooling conditions, the influence of different cooling speeds is here disregarded. The areas 1 to 4 in this diagram limit the chemical compositions of steels, where specific defects may occur during welding. Depending on the composition, purely ferritic chromium steels have a tendency to embrittlement by martensite and therefore to hot cracking (area 2) or to embrittlement due to strong
Nickel-equivalent = %Ni + 30x%C + 0,5x%Mn
grain growth (area 1). A cause for this strong grain growth during welding is the greatly increased diffusion speed in the ferrite compared with austenite. After reaching
a
temperature,
diffusion-start Figure
6.10
30
rit
28 26
0%
24
austenite
r Fe
5% 10
22
% %
0 A+F 2
20
40%
18 16
A +M
14
80%
12 10 8
100%
2
martensite F + M
00
2
6 4
4
6
A+M+F M+F ferrite 8
10
12 14 16
18
20 22 24
26 28
30 32
34
36 38
40
Chromium-equivalent = %Cr + %Mo + 1,5x%Si + 0,5x%Nb
shows that ferritic steels have
a
hardening crack susceptibility (preheating to 400°C!) hot cracking susceptibility above 1250°C
considerably
grain growth above 1150°C © ISF 2002
br-er-06-09e.cdr
stronger grain growth than
Schaeffler Diagram With Border Lines of Weld Metal Properties to Bystram
austenites. Therefore high alloyed ferritic steels are to
sigma embrittlement between 500-900°C
Figure 6.9
be considered as of limited weldability.
6000 m²
The area 3 marks a possible
5000
embrittlement of the material due to the development of σ-phase. As explained in 6.6, this risk occurs with increased increased
ferrite
contents,
chromium
grain size
4000
3000
2000
1000 ferritic steel
con-
tents, and sufficiently slow
austenitic steel
0
200
400
600
800
1000
°C
temperature
cooling speed.
br-er-06-10e.cdr
© ISF 2002
Grain Size as a Function of Temperature
Figure 6.10
1200
6. Welding High Alloy Steels
76
Finally, area 4 marks the strongly increased tendency to hot cracking in the austenite. Reason is, that critical elements responsible for hot cracking like e.g. sulphur and phosphorous have only very limited solubility in the austenite. During welding, they enrich the melt residue, promoting hot crack formation (see also chapter 9 - Welding Defects). There is a Z-shaped area in the centre of the diagram which does not belong to any other endangered area. This area of chemical composition represents the minimum risk of welding defects, therefore such a composition should be adjusted in the weld metal. Especially when welding austenitic steels one tries to aim at a low content of δ-ferrite, because it has a much greater solubility of S and P, thus minimising the risk of hot cracking. The Schaeffler diagram is not only used for determining the microstructure with known chemical composition. It is also possible to estimate the developing microstructures when welding different materials with or without filler metal. Figures 6.11 and 6.12 show two examples for a determination of the weld metal microstructures of so-called 'black and white' joints.
28 28
24 24
20
A+M
16
40 M
12
· 20% 123
² : ·=1:1
80
·
+
8
A+F
²
8
²
S235JR (St 37)
·
Welding consumable
12 16 20 24 Chromium-equivalent
80
2 1 30%
100 %
A+M+F M+F
+
F ·
F 0
28
32
· X10CrNiTi18-9 (W.-No. 1.4541) 21% Cr, 14% Ni, 3% Mo
4
8
36
12
16
20
24
28
32
Chromium-equivalent
²
S235JR (St 37)
·
Welding consumable
9
8
3 A+F
0 4
²
9
12
4
0 0
40
· M
F
M+F
F
16
² : ·=1:1
100 %
A+M+F
20
A+M
9
Nickel-equivalent
20
9
Nickel-equivalent
20
4
10
A 10
A
· X8Cr17 (W.-Nr. 1.4510) 21% Cr, 14% Ni, 3% Mo
Weld metal under 30 % dilution (= base metal amount)
Weld metal under 30 % dilution (= base metal amount)
© ISF 2002
br-er06-12e.cdr
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Application Example of Schaeffler - Diagram
Figure 6.12
© ISF 2002
Application Example of Schaeffler - Diagram
Figure 6.11
36
6. Welding High Alloy Steels
77
The ferrite content can only be measured with a relatively large dispersal, therefore DeLong proposed to base a measurement procedure on standardized specimens. Such a system makes it possible to measure comparable values which don't have to match the real ferrite content. Based on these measurement values, the ferrite content is no longer given in percentage, but steels are grouped by ferrite numbers. In addition to ferrite numbers, DeLong proposed a reworked Schaeffler diagram where the ferrite number can be determined by the chemical composition, Figure 6.13. Moreover, DeLong has considered the influence of nitrogen as a strong austenite developer (effects are comparable with influence of carbon). Later on, nitrogen was included into the nickel-equivalent of the Schaeffler diagram. Nickel-equivalent = %Ni + 30 x %C + 30 x %N + 0,5 x %Mn
21
e rit
19
r fe
nu
austenite
18
The most important feature
r be m 0 2
20
of high alloy steels is their 4
6
d re su ea ym all .-% tic vol e n in ag s m nt 0% ly te er con 2% rm fo rrite 4% Sc e f ha effl 6%% er6 au 7, 2% ste nite 9, 7% , -m art 10 ,3% en site 12 ,8% -lin 13 e
17 16 15 14 13 12 11 10 16
corrosion resistance start-
8 10 12 14 16 18
ing with a Cr content of 12%. In addition to the problems during welding
austenite + ferrite
described by the Schaeffler diagram, these steels can
17
18
26
25 19 20 21 22 23 24 Chromium-equivalent = %Cr + %Mo + 1,5 x %Si + 0,5 x %Nb
27
© ISF 2002
br-er-06-13e.cdr
be negatively affected with view to their corrosion re-
De Long Diagram
sistance caused by the
Figure 6.13 welding process.
Figure
air O
6.14 shows schematically
2Fe+++O+H2O ® 2Fe++++2OH-
the processes of electro-
OH
-
Fe+++
lytic
corrosion
under
water
a
drop of water on a piece of iron. In such a system a
O2
OH
H2O
opment of a local element
2Fe++
cathode anode
4e-
potential difference is a precondition for the devel-
Fe(OH)3
2Fe ® 2Fe+++4e-
O2+2H2O+4e ® 4OH
iron
-
© ISF 2002
br-er-06-14e.cdr
consisting of an anode and
Corrosion Under a Drop of Water
a cathode. To develop
Figure 6.14
6. Welding High Alloy Steels
78
such a local element, a different orientation of grains in the steel is sufficient. If a potential difference under a drop of water is present, the chemically less noble part reacts as an anode, i.e. iron is oxidised here and is dissolved as Fe2+-ion together with an electron emission. Caused by oxygen access through the air, a further oxidation to Fe3+ takes place. The cathodic, chemically nobler area develops OH- ions, absorbing oxygen and the electrons. Fe3+and OH--ions compose into the water-insoluble Fe(OH)3 which deposits as rust on the surface (note: the processes here described should serve as a principal explanation of electrochemical corrosion mechanisms, they are, at best, a fraction of all possible reactions). If the steel is passivated by chromium, the corrosion protection is provided by the development of a very thin chromium oxide layer which separates the material from the corrosive medium. Mechanical surface damages of this layer are completely cured in a very short time.
passive layer
active dissolution
passive layer
gap tensile stress
active dissolution of the crack base pitting corrosion passive layer
stress corrosion cracking passive layer
activly dissolved grain boundary chromium depleted zones active dissolution of the gap
crevice corrosion
grain boundary carbides
intergranular corrosion
incorrect
br-er06-15e.cdr
Figure 6.15
© ISF 2002
br-er06-16e.cdr
correct
© ISF 2002
Figure 6.16
The examples in Figure 6.15 are more critical, since a complete recovery of the passive layer is not possible from various reasons.
6. Welding High Alloy Steels
79
If crevice corrosion is present, corrosion products built up in the root of the gap and oxygen has no access to restore the passive layer. Thus narrow gaps where the corrosive medium can accumulate are to be avoided by introducing a suitable design, Figure 6.16.
br-er-06-17e.cdr
Pitting Corrosion of a Steel Storage Container
With pitting corrosion, the
Figure 6.17
chemical composition of the attacking medium causes a
local break-up of the passive layer. Especially salts, preferably Cl—ions, show this behaviour. This local attack causes a dissolution of the material on the damaged points, a depression develops. Corrosion products accumulate in this depression, and the access of oxygen to the bottom of the hole is obstructed. However, oxygen is required to develop the passive layer, therefore this layer cannot be completely cured and pitting occurs, Figure 6.17. Stress-corrosion cracking occurs when the material displaces under stress and the passive layer tears, Figure 6.18. Now the unprotected area is subjected to corrosion, metal is dissolved and the passive layer redevelops (figures 13). The repeated displace1
2
3
4
5
6
ment
and
repassivation
causes a crack propagation. 7
8
9
offset;
passive layer;
10
11
metal surface;
dislocation
12
Stress
cracking
corrosion
takes
mainly
place in chloride solutions. The crack propagation is transglobular, i.e. it does
br-er-06-18e.cdr
Model of Crack Propagation Through Stress Corrosion Cracking
Figure 6.18
not
follow
boundaries.
the
grain
6. Welding High Alloy Steels
80
Figure 6.19 shows the expansion-rate dependence of stress corrosion cracking. With very low expansion-rates, a curing of the passive layer is fast enough to arrest the crack. With very high expansion-rates, the failure of the specimen originates from a ductile fracture. In the intermediate range, the material damage is due to stress corrosion cracking. Figure 6.20 shows an example of crack propagation at transglobular stress corrosion cracking. A crack propagation speed is between 0,05 to 1 mm/h for steels with 18 - 20% Cr and 8 20% Ni. With view to welding it is important to know that already residual welding stresses
Sensitivity to stress corrosion cracking
may release stress corrosion cracking.
complete cover layer
tough fracture
T=RT
SpRK
e·2
e·1
Elongation speed e
·
© ISF 2002
br-er06-19E.cdr
br-er06-20e.cdr
Transgranular Stress Corrosion Cracking
Influence of Elongation Speed on Sensitivity to Stress Corrosion Cracking
Figure 6.19
© ISF 2002
Figure 6.20
The most important problem in the field of welding is intergranular corrosion (IC). It is caused by precipitation of chromium carbides on grain boundaries. Although a high solubility of carbon in the austenite can be expected, see Fe-C diagram, the carbon content in high alloyed Cr-Ni steels is limited to approximately 0,02% at room temperature, Figure 6.21.
6. Welding High Alloy Steels
81
The reason is the very high affinity of chromium to carbon, which causes the precipita-
to Bain and Aborn
Heat treatment temperature
1200
tion of chromium carbides Cr23C6 on grain
°C 1100
boundaries, Figure 6.22. Due to these precipitations, the austenite grid is depleted of
1000
chromium content along the grain boundaries A
900
and the Cr content drops below the parting limit. The diffusion speed of chromium in aus-
800
tenite is considerably lower than that of car700
bon, therefore the chromium reduction cannot
600 0
0.05
0.1 0.15 0.2 Carbon content
0.25 % 0,3
be compensated by late diffusion. In the depleted areas along the grain boundaries (line 2 in Figure 6.22) the steel has become susceptible to corrosion.
br-er06-21e.cdr
© ISF 2002
Carbon Solubility of Austenitic Cr - Ni Steels
Only after the steel has been subjected to sufficiently long heat treatment, chromium will
Figure 6.21
diffuse to the grain boundary and increase the
C concentration along the 1 - homogenuous starting condition 2 - start of carbide formation 3 - start of concentration balance 4 - regeneration of resistance limit
grain boundary (line 3 in Figure 6.22). In this way, the corrosion
resis-
tance can be restored (line 4 in Figure 6.22). Figure 6.23 explains why the IC is also described as intergranular
2 4
Chromium content of austenite
complete
1
resistance limit
disintegration. br-er-06-22e.cdr
3
Distance from grain boundary
Due to dissolution of deSensibility of a Cr - Steel
pleted areas along the grain boundary, complete grains break-out of the steel.
Figure 6.22
© ISF 2002
6. Welding High Alloy Steels
82
The precipitation and repassivation
mechanisms
described in Figure 6.22 are covered by intergranular corrosion diagrams according to Figure 6.24. Above a certain temperature carbon remains dissolved in the austenite © ISF 2002
br-er-06-23e.cdr
(see also Figure 6.21).
Grain Disintegration
Below this temperature, a carbon precipitation takes place. As it is a diffusion controlled
process,
Figure 6.23
the
precipitation occurs after a incubation
time
which depends on temperature (line 1, precipitation characteristic curve). During stoppage at a constant
temperature,
the
3 ¬ Reciprocal of heat treatment temperature 1/T
certain
unsaturated austenite
2
austenite chromium carbide (M23C6) no intergranular disintegration
austenite + chromium caride (M23C6) sensitive to intergranular disintegration
oversaturated austenite
1
parting limit of the steel is Heat treatment time (lgt)
regained by diffusion of chromium.
br-er-06-24e.cdr
1 incubation time 2 regeneration of resistance limit 3 saturation limit for chromium carbide
© ISF 2002
Area of Intergranular Disintegration of Unstabilized Cr - Steels
Figure 6.24 Figure 6.25 depicts characteristic precipitation curves of a ferritic and of an austenitic steel. Due to the highly increased diffusion speed of carbon in ferrite, shifts the curve of carbon precipitation of this steel markedly towards shorter time. Consequently the danger of intergranular corrosion is significantly higher with ferritic steel than with austenite.
6. Welding High Alloy Steels
83
As carbon is the element that triggers the intergranular corrosion, the intergranular corrosion diagram is relevantly influenced by the c content, Figure 6.26. By decreasing the carbon content of steel, the start of carbide precipitation and/or the start of intergranular corrosion are shifted towards
lower temperatures
and
longer
quench temperature
times. This fact initiated the development of
precipitation curves for 17% Cr steel
ELC-steels
(Extra-Low-Carbon)
18-8-Cr-Ni steel
Tempering temperature
so-called
where the C content is decreased to less than 0,03% During welding, the considerable influence of
cooling curve
carbon is also important for the selection of the shielding gas, Figure 6.27. The higher the CO2-content
of
the
shielding
gas,
Tempering time
the © ISF 2002
br-er06-25e.cdr
stronger is its carburising effect. The C-
Precipitation Curves of Various Alloyed Cr Steels
content of the weld metal increases and the steel becomes more susceptible to inter-
Figure 6.25
granular corrosion. An often used method to
1000 °C 900
avoid intergranular corro-
800
sion is a stabilisation of the steel by alloy elements like
700
Temperature
0.07%C
0.05%C 0.03%C
niobium and titanium, Fig-
600
ure 6.28. The affinity of
0.025%C
these elements to carbon is
500
significantly
higher
than
that of chromium, therefore 400 1 10 br-er-06-26e.cdr
10
2
104
103
105
Time
Influence of C-Content on Intergranular Disintegration
s
10
6
© ISF 2002
carbon is compounded into Nb- and Ti-carbides. Now carbon cannot cause any
Figure 6.26
chromium depletion. The
6. Welding High Alloy Steels
84
proportion of these alloy elements depend on the carbon content and is at least 5 times higher with titanium and 10 times higher with niobium than that of carbon. Figure 6.28 shows the effects of a stabilisation in the intergranular corrosion diagram. If both steels are subjected to the same heat treatment (1050°C/W means heating to 1050°C and subsequent water quenching), then the area of intergranular corrosion will shift due to stabilisation to significantly longer times. Only with a much higher heat treatment temperature the intergranular corrosion accelerates again. The cause is the dissolution of titanium carbides at sufficiently high temperature. This carbide dissolution causes problems when welding stabilised steels. During welding, a narrow area of the HAZ is heated above 1300°C, carbides are dissolved. During the subsequent cooling and the high cooling rate, the carbon remains dissolved.
0.058 % C 0.53 % Nb Nb/C = 9
°C 600
0.030 % C 0.51 % Nb Nb/C = 17
0.018 % C 0.57 % Nb Nb/C = 32
M2
550
M1 500
S1
450
Heat treatment temperature
Heat treatment temperature
700 800 °C 700
1050°C
650
/W
600 550 500 450 0,3
400 0,2
0,5
1
2,5
5
10
50
25
100
250
h
1000
Heat treatment time
1 3 10 30 100 Time W.-No.:4301 (0,06%)
Heat treatment temperature
X5CrNi18-10
C o m p o sitio n S hie ld ing g a s
A r [% ]
C O2
O2
S 1
99
/
1
M 1
90
5
5
M 2
82
18
/
Influence of Shielding Gas on Intergranular Disintegration
h
10000
unstabilized
650
1300°C
/W
600 1050°C
550
/W
500 450 0,3
© ISF 2002
br-er06-27e.cdr
1000
800 °C 700
1 3 W.-No.:4541
X5CrNiTi18-10
Figure 6.27
300
10
30
100 Time
300
1000
h
10000
stabilized
© ISF 2002
br-er06-28e.cdr
Influence of Stabilization on Intergranular Disintegration
Figure 6.28
If a subsequent stress relief treatment around 600°C is carried out, carbide precipitations on grain boundaries take place again. Due to the large surplus of chromium compared with niobium or titanium, a partial chromium carbide precipitation takes place, causing again inter-
6. Welding High Alloy Steels
85
granular susceptibility. As this susceptibility is limited to very narrow areas along the welded joint, it was called knife-line attack because of its appearance. Figure 6.29. In stabilised steels, the chromium carbide represents an unstable phase, and with a sufficiently long heat treatment to transform to NbC, the steel becomes stable again. The stronger the steel is over-stabilised, the lower is the tendency to knife-line corrosion. Nowadays the importance of Nickel-Base-Alloys increases constantly. They are ideal materials when it comes
to
components
which are exposed to special conditions: high temperature, corrosive attack, low temperature, wear rebr-er-06-29e.cdr
sistance, or combinations
Knife-Line Corrosion
hereof. Figure 6.30 shows one of the possible group-
Figure 6.29
ing of nickel-base-alloys. Materials listed there are selected examples, the total number of available materials is many times higher. Group A consists of nickel alloys. These alloys are Alloy
Chem. composition
Alloy
Nickel 200
Ni 99.6, C 0.08
Duranickel 301 Ni 94.0, Al 4.4, W 0.6
Nickel 212 Nickel 222
Ni 97.0, C 0.05, Mn 2.0 Ni 99.5, Mg 0.075
Incoloy 925 Ni 42.0, Fe 32.0, Cr 21.0, Mo 3.0, W 2.1, Cu 2.2, Al 0.3 Ni-Span-C 902 Y2O3 0.5, Ni 42.5, Fe 49.0, Cr 5.3, W 2.4, Al 0.5
Group A
Chem. Composition
characterized by moderate
Group D1
Group B
Group D2
Monel 400
Ni 66.5, Cu 31.5
Monel K-500
Ni 65.5, Cu 29.5, Al 2.7, Fe 1.0, W 0.6
Monel 450 Ferry Group C
Ni 30.0, Cu 68.0, Fe 0.7, Mn 0.7
Inconel 718
Ni 52.0, Cr 22.0, Mo 9.0, Co 12.5, Fe 1.5, Al 1.2
Ni 45.0, Cu 55.0
Inconel X-750 Ni 61.0, Cr 21.5, Mo 9.0, Nb 3.6, Fe 2.5 Nimonic 90 Ni 77.5, Cr 20.0, Fe 1.0, W 0.5, Al 0.3, Y2O3 0.6
Inconel 600
Ni 76.0, Cr 15.5, Fe 8.0
Nimonic 105
Ni 76.0, Cr 19.5, Fe 112.4, Al 1.4
Nimonic 75
Ni 80.0, Cr 19.5
Incoloy 903
Ni 39.0, Fe 34.0, Cr 18.0, Mo 5.2, W 2.3, Al 0.8
Nimonic 86
Ni 64.0, Cr 25.0, Mo 10.0, Ce 0.03
Incoloy 909
Ni 58.0, Cr 19.5, Co 13.5, Mo 4.25, W 3.0, Al 1.4
Incoloy 800
Ni 32.5, Fe 46.0, Cr 21.0, C 0.05
Inco G-3
Ni 38.4, Fe 42.0, Cu 13.0, Nb 4.7, W 1.5, Al 0.03, Si 0.15
Incoloy 825
Ni 42.0, Fe 30.0, Cr 21.5, Mo 3.0, Cu 2.2, Ti 1.0
Inco C-276
Ni 38.4, Fe 42.0, Cu 13.0, Nb 4.7, W 1.5, Al 0.03, Si 0.4
Inco 330
Ni 35.5, Fe 44.0, Cr 18.5, Si 1.1
Group E Monel R-405
mechanical strength and high degree of toughness. They can be hardened only by cold working. The alloys are quite gummy in the annealed or hot-worked con-
Ni 66.5, Cu 31.5, Fe 1.2, Mn 1.1, S 0.04
dition,
and
cold-drawn
© ISF 2002
br-er-06-30e.cdr
material is recommended Typical Classification of Ni-Base Alloys
Figure 6.30
for best machinability and smoothest finish.
6. Welding High Alloy Steels
86
Group B consists mainly of those nickel-copper alloys that can be hardened only by cold working. The alloys in this group have higher strength and slightly lower toughness than those in Group A. Cold-drawn or cold-drawn and stress-relieved material is recommended for best machinability and smoothest finish. Group C consists largely of nickel-chromium and nickel-iron-chromium alloys. These alloys are quite similar to the austenitic stainless steels. They can be hardened only by cold working and are machined most readily in the cold-drawn or cold-drawn and stress-relieved condition. Group D consists primary of age-hardening alloys. It is divided into two subgroups: D 1 – Alloys in the non-aged condition. D 2 – Aged Group D-1 alloys plus several other alloys in all conditions. The alloys in Group D are characterized by high strength and hardness, particularly when aged. Material which has been solution annealed and quenched or rapidly air cooled is in the softest condition and does machine easily. Because of softness, the non-aged condition is necessary for trouble free drilling, tapping and all threading operations. Heavy machining of the age-hardening alloys is best accomplished when they are in one of the following conditions: 1. Solution annealed 2. Hot worked and quenched or rapidly air cooled Group E contains only one material: MONEL R-405. It was designed for mass production of automatically machined screws. Due to the high number of possible alloys with different properties, only one typical material of group D2 is discussed here: Material No. 2.4669, also known as e.g. Inconel X-750. The aluminium and titanium containing 2.4669 is age-hardening through the combination of these elements with nickel during heat treatment: gamma-primary-phase (γ') develops which is the intermetallic compound Ni3(Al, Ti). During solution heat treatment of X-750 at 1150°C, the number of flaws and dislocations in the crystal is reduced and soluble carbides dissolve. To achieve best results, the material
6. Welding High Alloy Steels
87
should be in intensely worked condition before heat treatment to permit a fast and complete recrystallisation. After solution heat treatment, the material should not be cold worked, since this would generate new dislocations and affect negatively the fracture properties. The creep rupture resistance of X-750 is due to an even distribution of the intercrystalline γ' phase. However, fracture properties depend more on the microstructure of the grain boundaries. During an 840°C stabilising heat treatment as part of the triple-heat treatment, the fine γ' phase develops inside the grains and M23C6 precipitates onto the grain boundaries. Adjacent to the grain boundary, there is a γ' depleted zone. During precipitation hardening (700°C/20 h) γ' phase develops in these depleted zones. γ' particles arrest the movement of dislocations, this leads to improved strength and creep resistance properties. During the M23C6 transformation, carbon is stabilised to a high degree without leaving chromium depleted areas along the grain boundaries. This stabilisation improves the resistance of this alloy against the attack of several corrosive media. With a reduction of the precipitation temperature from 730 to 620°C – as required for some special heat treatments – additional γ' phase is precipitated in smaller particles. This enhances the hardening effect and improves strength characteristics. Further metallurgical discussions about X-750, can be taken from literature, especially with view to the influence of heat treatment on fracture properties and corrosion behaviour.
The recommended processes for welding of X-750 are tungsten inert gas, plasma arc, electron beam, resistance, and pressure oxy arc welding. During TIG welding of INCONEL X-750, INCONEL 718 is used as welding consumable. Joint properties are almost 100% of base material at room temperature and about 80% at 700° 820°C. Figure 6.31 shows typical strength properties of a welded plate at a temperature range between -423° and 1500°F (-248 – 820°C). Before welding, X-750 should be in normalised or solution heat treated condition. However, it is possible to weld it in a precipitation hardened condition, but after that neither the seam nor the heat affected zone should be precipitation hardened or used in the temperature range of precipitation hardening, because the base material may crack. If X-750 was precipitation hardened and then welded, and if it is likely that the workpiece is used in the temperature range of precipitation hardening, the weld should be normalised or once again precipitation hardened. In any case it must be noted that heat stresses are minimised during assembly or welding.
6. Welding High Alloy Steels
88
X-750 welds should be solution heat treated before a precipitation hardening. Heating-up speed during welding must be from the start fast and even touching the temperature range of precipitation hardening only as briefly as possible. The best way for fast heating-up is to insert the welded workpiece into a preheated furnace. Sometimes a preheating before welding is advantageous – if the component to be welded has a poor accessibility, or the welding is complex, and especially if the assembly proves to be too complicated for a post heat treatment. Two effective welding preparations are: 1. 1550°F/16 h, air cooling 2. 1950°F/1 h, furnace cooling with 25°-100°F/h up to 1200°F, air A repair welding of already fitted parts should be followed by a solution heat treatment (with a fast heating-up through the temperature range of precipitation hardening) and a repeated precipitation hardening. A cleaning of intermediate layers must be 220
are formed during welding. (A complete isola-
200
tion of the weld metal using gas shielded
180
processes is hardly possible). If such films are not removed on a regular basis, they can
Stress, 1000 psi
carried out to remove the oxide layers which
160 tensile strength 140
become thick enough to cause material sepa-
120
rations together with a reduced strength.
100 0.2% yield stress 80
Brushing with wire brushes only polishes the blasted or ground with abrasive material. The
Elongation, %
60
surface, the layer surface must be sand-
30 20
elongation in 1/2”
10 0
frequency of cleaning depends on the mass
20
of the developed oxides. Any sand must be
0
elongation in 2”
10 -423
0
800
1000
1200
1400
1600
© ISF 2002
Mechanical Properties of a Typical Ni-Base Alloy
equipment must be of adequate performance. malized or solution heat treated condition.
600
br-er06-31e.cdr
, seam-, and flash butt welding. The welding X-750 is generally resistance welded in nor-
400
Temperature, F°
removed before the next layer is welded. X-750 can be joined also by spot-, projection-
200
Figure 6.31
7. Welding of Cast Materials
7. Welding of Cast Materials
90
Figure 7.1 pro-
cast materials
vides a summary of
the
metallic cast materials
non-metallic cast materials plastics, gypsum and s.th.similar
different
iron-carboncast materials
non-iron-metal cast materials
cast iron materials. In this connection it is only
unalloyed ferritic
referred iron,
to
cast
cast
and
nodular graphite cast iron
ferritic
perlitic
steel
lemellar graphite cast iron
high alloyed
hard cast clear chill low C iron casting content
high Ccontent
austenitic
ferritic not decarburized
decarburized
decarburized annealed malleable cast iron
malleable
cast
low alloyed
alloyed
perlitic
steel, as special
special cast iron (G...)
cast iron
malleable iron
cast steel
ferritic
perlitic
Cr-cast iron
not decarburized annealed malleable cast iron
perlitic
ferritic
perlitic
ledeburitic
austenitic
graphite
other elements
Si-cast iron
Al-cast iron
austenitic © ISF 2002
br-er-07-01e.cdr
materials,
Table of the cast Iron Materials
due to their poor weldability, are of
Figure 7.1
no importance in welding. Figure 7.2 shows the designation of Designation according to the material code (DIN EN 1560)
the cast material in accordance with
e.g.: EN-GJ L F – 150
DIN EN 1560. A distinction is made 1 Position 1: Position 2: Position 3: Position 4: Position 5:
EN GJ L F 150
Position 6:
-
2 34
5
between the designation “according to
standardised material cast material graphite structure (lamellar graphite) microstructure (ferritic) 2 mechanical properties (Rm= 150 N/mm ) chemical composition (high alloyed) optionally
the material code” and the designation “according to the material number”. In Figure 7.2, examples of two materials are specified.
Designation according to the material number
e.g.: EN- J L 1271 1 23 Position 1: Position 2: Position 3: Position 4: Position 5: Position 6:
EN J L 1 27 1
-
4,5,6
standardised material cast material graphite structure (lamellar graphite) number for the main characteristic material identification number special requirement
© ISF 2004
br-er07-02e.cdr
Designation of Materials
Figure 7.2
7. Welding of Cast Materials
91
Figure 7.3 depicts a survey of the mechanical properties and the chemical compositions of several customary cast materials. As to its analysis and mechanical properties which are very different from other cast materials, cast steel constitutes an exception to the rule.
In Figure 7.4 the stable and the metastable iron-carbon diagram are shown. The differences between the cast material are best explained this way. Cast iron with lamellar and spheroidal
graph-
ite
carbon
has
contents tween
of 2,8
beand
4,5%. Through the addition of alloying elements,
above
all Si, these mateFigure 7.3
rials solidify fo llowing the stable system, i.e., the carbon is precipitated in
the
form
of
graphite. Malleable cast
iron
shows
similar C-contents, the
solidification
from
the
molten
metal,
however,
follows
the
tastable
me-
system.
The C-contents of cast steel, on the Figure 7.4
7. Welding of Cast Materials
92
other hand, comply with those of common structural steels, i.e., they are, as a rule, below 0,8% C.
The structure of a normalised cast iron which is composed of ferrite (bright) and pearlite (dark) is shown in Figure 7.5. Since the properties are similar to those of structural steels these materials are weldable, constructional welding is also possible. It is recommended to normalise the cast steel parts before welding. Through this type of heat treatment, on the one hand the transformation of the cast structure is obtained, the residual stresses inside the workpiece are, on the other hand, reduced.
Figure 7.5
From a C-content in the steel cast of 0,15% up, it is recommended to carry out preheating during welding, the preheating temperature should follow the analysis of the material, the workpiece geometry and the welding method. After welding the cast workpieces are subject to stress-relief annealing.
Figure 7.6 shows the structure of cast iron with lamellar graphite (grey cast iron). Apart from their carbon content, these materials are characterised by increased contents of S and P which
Figure 7.6
7. Welding of Cast Materials
93
improves castability. Besides the poor mechanical properties (elongation after fracture of approx. 1%), these chemical properties also impede welding with ordinary means. It is not possible to carry out constructional welding with grey cast iron. Repair welds of grey cast iron are, in contrast, carried out more frequently as damaged cast parts are not easily replaceable. For those repair welds, the cast parts must be preheated (entirely or partly) to te mperatures of approx. 650°C. Heating and cooling must be done very slowly as the cast piece may be destroyed already by the thermal stresses. The highly liquid weld metal also constitutes a problem, and thus the molten pool must be supported by a carbon pile. Welding may be carried out with similar filler material (materials of the same composition as the base). If grey cast iron is to be welded without any preheating, the filler material must, as a rule, be dissimilar (of different composition to the base metal). During this type of welding, there are always strong structural changes in the region of the weld which lead to high hardening and high residual stresses. For the minimisation of these structural changes, a highly ductile filler material is applied. The heat input into the base material should be as low as possible. Figure 7.7 depicts the structural constitution of sphe roidal graphite cast iron. The graphite spheroidization achieved
by
is the
addition of magnesium and cerium. As, with this type of
graphite,
the
notch actions are Figure 7.7
considerably lesser than this is
the case with grey cast iron, this type of cast iron is characterised by substa ntially better mechanical parameters with a considerably higher elongation after fracture and improved ductility. For this reason, the risk of material failure caused by weld residual stresses or thermal stresses is considerably reduced for spheroidal graphite
7. Welding of Cast Materials
94 cast iron. Frequently, nickel-based alloys are used as filler material. Problems occur in the HAZ where, besides the ledeburite eutectic alloy system, also Ni-Fe-martensite is frequently formed. Both structures lead to extreme hardening in the HAZ which can
be
removed
only
consuming heat treatment.
Figure 7.8
Figures 7.8 and 7.9 show the structures of Carburized Annealed Malleable Cast Iron (7.7) and of Decarburized Annealed Malleable Cast Iron (7.9). The composition of the malleable cast iron is thus that during solidification, the total of carbon is bound in cementite and precipitated. During a subsequent annealing process, the iron carbide disintegrates into graphite and iron.
Figure 7.9
by
time-
7. Welding of Cast Materials
95 If annealing is carried out in neutral atmosphere, the structure of Carburized Annealed Malleable Cast Iron develops. Annealing in oxidising atmosphere leads to the decarburisation of the workpiece surfaces and Decarburized Annealed Malleable Cast Iron is developed, Figure 7.10. Carburized
Annealed
Malleable
Cast Iron is not weldable. Decarburized Annealed Malleable Cast Iron, in contrast, is weldable.
Figure 7.10
You can see in Figure 7.11 that, also with malleable cast iron, hardening in the region of the HAZ occurs. For carrying out constructional welds made of malleable cast iron parts, a special material quality has been developed. Figure 7.11 shows that this material, EN-GJMW-400-12, is characterised by considerably less hardening. This material is weldable without any problems up to a wall thickness of 8 mm.
Figure 7.11
8. Welding of Aluminium
8. Welding of Aluminium
97 Figure 8.1 compares basic physical properties
Property
Al
Fe
of steel and aluminium. Side by side with different mechanical behaviour, the following
Atomic weight
[g/Mol]
26.9
55.84
Specific weight
[g/cm³]
2.7
7.87
fcc
bcc
Lattice E-module
[N/mm²]
71*10³
210*10³
R pO,2 PO,2
[N/mm²]
ca. 10
ca. 100
R mm
[N/mm²]
ca. 50
ca. 200
spec. Heat capacity
[J/(g*°C)]
0.88
0.53
[°C]
660
1539
[W/(cm*K)]
2.3
0.75
Spec. el. Resistance
[nWm]
28-29
97
Expansion coeff.
[1/°C]
24*10 -6
12*10 -6
Melting point Heat conductivity
FeO Oxydes
Al2O 3
Melting point of oxydes
[°C]
2050
Fe 3O 4
differences are important for aluminium welding: - considerably lower melting point compared with steel - three times higher heat conductivity - considerably lower electrical resistance - double expansion coefficient - melting point of Al203 considerably higher
Fe 2O 3
than that of Al; metal and iron oxide melt ap-
1400
proximately at the same temperature.
1600 (1455) © ISF 2002
br-er08-01.cdr
Basic Properties of Al and Fe
Figure 8.2 compares some mechanical properties of steel with properties of some light metals. The important advantages of light
Figure 8.1
metals compared with steel are especially
shown in the right part of the figure. If a comparison should be based on an identical stiffness, then the aluminium supporting beam has a 1.44 times larger cross-section than the steel beam, however only about 50% of its weight. Figure 8.3 compares qualitatively the stress-strain diagram
of
Aluminium
and
steel. In contrast to steel, aluminium has a fcc (face centred
cubic)-lattice
at
room temperature. This is why there is no distinct yield point as being the case in a bcc (body centred cubic)lattice.
Aluminium
is
br-er-08-02.cdr
Deflexions and Weights of Cantilever Beams Under Load
not
subject to a lattice trans-
Figure 8.2
8. Welding of Aluminium
98
formation during cooling, thus there is no structure transformation and consequently no danger of hardening in the heat affected zone as with steel.
4 cm 2
low carbon steel
200°C
400
1000 1200
600
800
1500
-2
Steel
-4
Stress
8 cm aluminium 6
100°C 200
4
Al-alloy
2 300 400 500
600
-2 -4 -6 -8 -18
Elongation © ISF 2002
br-er08-03.cdr
-16
-14
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Comparison of Stress-Elongation Diagrams of Al and Steel
Figure 8.3
-12
-10
-8
-6
-4
-2
0
2
cm
6
© ISF 2002
Isothermal Curves of Steel and Al
Figure 8.4
Figure 8.4 illustrates the effect of the considerably higher heat conductivity on the welding process compared with steel. With aluminium, the temperature gradient around the welding point is considerably smaller than with steel. Although the peak temperature during Al welding is about 900°C below steel, the isothermal curves around the welding point have a clearly larger extension. This is due to the considerably higher heat conductivity of aluminium compared with steel. This special characteristic of Al requires a input heat volume during welding equivalent to steel. Figure 8.5 lists the most important alloy elements and their combinations for industrial use. Due to their behaviour during heat treatment can Al-alloys be divided into the groups hardenable and non-hardenable (naturally hard) alloys.
8. Welding of Aluminium
99
Al Cu Mg
ing consumables. Al Mg Si
Cu
Aluminium alloys are often welded with con-
Al Zn Mg
sumable of the same type, however, quite Mg
often over-alloyed consumables are used to
Al Zn Mg Cu
678
Al alloys together with preferably used weld-
hardenable alloys
Figure 8.6 shows typical applications of some
Al
Zn
Al Si Cu
and to improve the mechanical properties of Al Si
the seam.
Si Al Mg
The classification of Al alloys into two groups
Al Mg Mn
Mn
is based on the characteristic that the group Al Mn
of the non-hardenable alloys cannot increase
© ISF 2002
br-er08-05.cdr
the strength through heat treatment, in con-
678
Mg and Zn because of their low boiling point)
non-hardenable alloys
compensate burn-off losses (especially with
Classification of Aluminium Alloys
trast to hardenable alloys which have such a potential. The important hardening mechanism for this
Figure 8.5
second group is explained by the figures 8.7 und 8.8. Example: If an alloy containing about 4.2% Cu, which is stable at room temperature, is heat treated at 500°C, then, after a sufficiently long time, there will be only a single phase structure present. All alloy elements were dissolved, Figure 8.8 between point P and Q. When quenched to room Al - alloys Al99,5 AlCuMg1 AlMgSi0,5 AlSi5 AlMg3
AlMg2Mn0,8 AlMn1
Typical use electrical engineering mechanical engineering, food industries architecture, electrical engineering, anodizing quality architecture, anodizing quality architecture, apparatus-, vehicle-, shipbuilding engineering, furniture industry apparatus-, vehicle-, shipbuilding engineering apparatus-, vehicle-engineering, food industry
W elding consumable SG-Al 99,5Ti; SG-Al 99,5
tion, no precipitation will
SG-AlMg4,5Mn
take place. The alloy ele-
SG-AlMg5; SG-AlMg4,5Mn; SG-AlSi5 SG-AlSi5
ments are forced to remain dissolved, the crystal is out
SG-AlMg3; SG-AlMg4,5Mn SG-AlMg5; SG-AlMg3; SG-AlMg4,5Mn
of equilibrium. If such a structure is subjected to an
SG-AlMn1;SG-Al99,5T
age hardening at room or
base material - aluminium percentage of alloy elements without factor
elevated
temperature,
a
© ISF 2002
br-er-08-06.cdr
Use and Welding Consumables of Aluminium Alloys
Figure 8.6
temperature in this condi-
precipitation of a second phase takes place in ac-
8. Welding of Aluminium
100
cordance with the binary system, the crystal tries to get back into thermodynamical equilibrium. Depending on the level of
stable condition
solution heat treatment
repeated hardening
solidification of alloy elements in solid solution
hardening temperature, the
quenching
regeneration
oversaturated solid solution, metastable condition
precipitation takes place in
warm ageing
cold ageing (RT ageing)
ageing at slightly increased temperature coherent precipitations, cold aged condition
three possible forms: copartly coherent precipitations, warm aged condition
coherent and partly coherent precipitations, transition conditions cold ageing -- warm ageing temperature rise
temperature rise
herent particles (i.e. particles
longer warm ageing partly coherent and incoherent precipitations, softening
from
the
matrix in their chemical composition but having the
longer warm ageing stable incoherent equilibrium phase stable condition © ISF 2002
br-er-08-07.cdr
deviating
Ageing Mechanism
same
lattice
structure),
partly
coherent
particles
(i.e. the lattice structure of the matrix is partly re-
Figure 8.7
tained),
and
incoherent
particles (lattice structure completely different from the matrix), Figure 8.7. Coherent particles formed at room temperature can be transformed into incoherent particles by increase of temperature (i.e. enabling diffusion). The precipitations cause a restriction to the
700 liquid
dislocation movement in the matrix lattice, thus
liquid and solid Q
600
leading to an increase in strength. The finer the
copper containing aluminium solid solution 500
At an increased temperature (heat ageing, Fig-
Temperature
precipitations, the stronger the effect.
P
400
300 aluminium solid solution and copper aluminide (Al2Cu)
ure 8.7) a maximum of second phase has precipitated after elapse of a certain time. Consequently a prolonged stop at this tem-
200
100 copper content of AlCuMg
perature does not lead to an increased strength, but to coarsening of particles due to
0
1
2
3
4
5
mass-%
Copper
diffusion processes and to a decrease in strength (less bigger particles in an extended
br-er08-08.cdr
space).
© ISF 2002
Phase Diagram Al-Cu
Figure 8.8
7
8. Welding of Aluminium
101 After a very long heat ageing a stable condition is reached again with relatively large precipitations of the second phase in the matrix. In Figure 8.7 is this stable final condition iden-
Q
tical with the starting condition. A deteriorati-
solution heat treatment
500 P
on of mechanical properties only happens
°C
quenching
Temperature
400
during hot ageing, if the ageing time is excessively long.
300
200
heat ageing
The complete process of hardening at room
100
temperature is metallographic also called age age hardening
hardening, at elevated temperature heat age0
2
4
6
8
10
12
h
14
Time
ing. A decrease in strength at too long ageing time is called over-ageing.
br-er08-09.cdr
© ISF 2002
Temperature - Time Distribution During Ageing
Figure 8.9 shows a schematic representation of time-temperature curves during hardening
Figure 8.9
Figure
with age hardening and heat ageing.
8.10
shows
the
380
strength increase of AlZnMg The difference between age hardening and heat ageing is here very clear. Due to improved
diffusion
condi-
tions is the strength increase
320 0.2% yield stress s0.2 in N/mm²
1 in dependence of time.
water quenching (~900°C/min) air cooling (~30°C/min)
260
120°C 200
RT 140
80 10-1
in the case of heat ageing much faster than in the case of
age
hardening.
quenched
100
101
10²
10³
Ageing time in h © ISF 2002
br-er-08-10.cdr
Increase of Yield Stress During Ageing of AlZnMg1
The
strength maximum is also reached considerably ear-
Figure 8.10
lier. The curve of hot ageing shows clearly the begin of strength loss when held at a too long stoppage time. This figure shows another specialty of the process of ageing. During ageing, a
8. Welding of Aluminium
102
second phase is precipitated from a single-phase structure. To initiate this process, the structure must contain nuclei of the second phase. However, a certain time is required to develop such nuclei. Only after formation of nuclei can the increase in strength start. The period up to this point is called incubation time. 500 110
N/mm² Tensile strength sB
Figure 8.11 shows the effect of the height of ageing temperature level on both, mechanical properties of a hardenable Al-alloy and on in-
135
400
150
180
300
190 205
230
260°C
cubation time. The lower the ageing tempera-
200 110
N/mm² 400 0.2% yield stress s0.2
ture, the higher the resulting values of yield stress and tensile strength. If a low ageing temperature is selected, the ageing time as well as
300
135
200
180 190 205°C
150
the incubation time become extremely long. Fracture elongation d2
230 260
Figure 8.11 shows that a the maximum yield stress is reached after a period of about one
190
180
205
150
135
20
10
110°C
260
230 1 day
30 min
10
0
year under a temperature of 110°C. An increase of the ageing temperature shortens the
30 %
-2
10
-1
0
1
10 10 Ageing time
1 week
10
2
1 1 month year
10
3
h 10
4
br-er08-11.cdr
© ISF 2002
Influence of Ageing Temperature and -Time on Ageing
duration of the complete precipitation process by a certain value raised by 1 to a power. On the other hand, such an acceleration of ageing
Figure 8.11 leads to a lowering of the
400
maximum strength. As the
N/mm²
lower part of the figure
Tensile strength Rm
300
AlMg5
shows, the fracture elonga-
AlMg3
tion is counter-proportional
200
to the strength values, i.e. the
100
strength
increase
caused by ageing is ac-
Al99,5
companied by an embrit0 0
30
%
70
Age Hardening of Al Alloys
Figure 8.12
Strain © ISF 2002
br-er-08-12.cdr
tlement of the material.
8. Welding of Aluminium
103
Figure 8.12 shows a method of how to increase the strength of non-hardenable alloys. As no precipitations are present to reduce the movement of dislocations, such alloys can only be strengthened by cold working. Figure 8.12 illustrates two essential mechanisms of strength increase of such alloys. On 300
one hand, tensile strength increases with in-
N/mm²
creasing content of alloy elements (solid solu-
250
tion strengthening), on the other hand, this increase is caused by a stronger deformation
Rm or Rp0,2
200
of the lattice. 150
Figure 8.13 shows the effect of the welding process on mechanical properties of a cold-
0,7
100
worked alloy. Due to the heat input during
0,5 50 HV30
0,4
Rp0,2/Rm
0,6
(recovery), in addition, a grain coarsening will
0,3 0,2
0 80
60 40 20 0 20 40 Distance from Seam Centre
welding, the blocked dislocations are released start in the HAZ. This is followed by a strong
60 mm 100
drop in yield point and tensile strength. This
br-er08-13.cdr
strength loss cannot be overcome in the case
© ISF 2002
Non-Hardenable Al Alloy
of a welding process.
Figure 8.13 400
Figure
8.14
illustrates
the
90 days RT
N/mm²
Rm
350
mechanisms in the case of a
21 days RT
hardenable aluminium alloy. welding heat, the precipitations are solution heat treated
Rp0,2
250
90 days RT
Stress
As a consequence of the
1 day RT
300
21 days RT 200
4 mm plates of: AlZnMg1F32 start values: Rp0,2=263N/mm² Rm=363 N/mm² welding method: WIG, both sides, simultaneously welding consumable: S-AlMg5 specimens with machined weld bead
1 day RT
150
and the strength values de100
crease in the weld area. Due to the age hardening, a re-
50 80 br-er-08-14.cdr
strengthening of the alloys
40
20
20 60 0 40 Distance from seam centre
Hardenable Al Alloy
takes place with increasing time.
60
Figure 8.14
80
100
mm
140 © ISF 2002
8. Welding of Aluminium
104 Figure 8.15 shows another problematic nature of Alwelding. Due to the high thermal expansion of aluminium, high tensions develop during solidification of the weld pool in the course of the welding cycle. If the welded alloy indicates
© ISF 2002
br-er-08-15.cdr
a
high
melting
interval, cracks may easily
Hot Cracks in a Al Weld
develop in the weld. Figure 8.15 A relief can be afforded by preheating of the material, Figure 8.16. With an increasing preheat temperature, the amount of fractured welds decreases. The different behaviour of the three displayed alloys can be explained using the right part of the figure. One can
100 %
see that the manganese signifi-
cantly the hot crack susceptibility. The maximum of this
2 60 1 40
X X
3
20
hot crack susceptibility is
Mg
Cracking susceptibility
influences
Weld cracking tendency
content
80
Si
X X
likely with about 1% Mg content (corresponds with alloy 1). With increasing MG con-
0
100
300
Preheat temperature
400
°C
500
0
1
2
3
%
4
Alloy content 1: AlMgMn 2: AlMg 2,5 3: AlMg 3,5
© ISF 2002
br-er-08-16.cdr
tent, hot crack susceptibility
Influence of Preheat Temperature and Magnesium Content
decreases strongly (see also alloy 2 and 3, left part).
200
Figure 8.16
To avoid hot cracking, partly very different preheat temperatures are recommended for the alloys. Zschötge proposed a calculation method which compares the heat conductivity conditions of the Al alloy with those of a carbon steel with 0.2% C. The formula is shown in Figure
8. Welding of Aluminium
105
8.17, together with the related calculation result. These results are only to be regarded as approximate, the individual application is subject to the information of the manufacturer.
strong
porosity
of
the
TS Tvorw. lAl-Leg.
in °C in °C in J/cm*s*K 660
the interplay of several characteristics and hard to suppress. Pores in Al are mostly formed
by
hydrogen,
temperature of melt start (solidus temperature) preheat temperature heat conductivity
melting point pure aluminium
600 Recommended preheat temperature
welded joint. It is based on
745 l Al-Leg.;
°C 500 400 300 200
Welding possible without preheating: AlMg5, AlMg7, AlMg4.5Mn, AlZnMg3, AlZnMg1
100
0
which is driven out of the
mild steel (0.2%C) without preheating
during Al welding is the
TVorw. = TS -
Al Zn Mg Cu 0,5 Al Zn Mg Cu 1,5
problem
Al Si 5 Al Cu Mg 1 Al R Mg 2 Al Cu Mg 0,5 Al Mn Al Mg 2 Al Cu Mg 2 Al Mg 3 Al Mg 3 Si Al Mg Mn
major
Al 99,98R Al99,9 Al99,8 Al 99,7 Al 99,5 Al 99 Al R Mg0,5 Al Mg Si 0,5 Al Mg Si 0,8 Al Mg Si 1 E Al Mg Si 1 Al Mg 1
Another
Increasing better weldability © ISF 2002
br-er-08-17.cdr
weld pool during solidifica-
Recommendations for Preheating
tion. Solubility of hydrogen in
aluminium
abruptly
on
changes the
Figure 8.17
phase
transition melt-crystal, i.e. the melt dissolves many times more of the hydrogen than the just forming crystal at the same temperature. This leads to a surplus of hydrogen in the melt due to the crystallisation during solidification. This surplus precipitates in
Excessive Porosity in a Al Weld
form of a gas bubble at the solidifying front. As the
© ISF 2002
br-er-08-18.cdr
Figure 8.18
melting point of Al is very low and Al has a very high heat conductivity, the solidification speed of Al is relatively high. As a result, in the melt ousted gas bubbles have often no chance to rise all the way to the surface. Instead, they are passed by the solidifying front and remain in the weld metal as pores, Figure 8.18.
8. Welding of Aluminium
106 To suppress such pore
irregular wire electrode feed
too thick and water containing oxyde layer by too long or open storage in non air-conditioned rooms
formation it is therefore
humid air (nitrogen, oxygen, water)
necessary to minimise the
nozzle deposits and too steep inclination of the torch cause turbulences
poor current transition
VS
humid air
hydrogen content in the melt. Figure 8.19 shows
too thick oxyde layer (condensed water) dirt film (oil, grease)
possible sources of hydrogen during MIG welding of
feuchte Luftpores Poren solid weld metal
base material
Al.
H2 H2
festes Schweißgut Grundwerkstoff
© ISF 2002
br-er-08-19.cdr
Ingress of Hydrogen Into the Weld
Figure 8.19
parallel gap
Figure 8.20 and 8.21 show the effect of pure thermal
weld pool
expansion during Al welding. The large thermal expansion
overlap
of the aluminium along with the relatively large heat affected zones cause in com-
opening gap
bination with a parallel gap adjustment a strong distor-
weld pool
tion of the welded parts. To minimise this distortion, the workpieces must be set at a
© ISF 2002
br-er-08-20.cdr
suitable angle before weld-
Weld Gap Adjustment
ing, Figure 8.21. Figure 8.20
8. Welding of Aluminium
wedge
br-er08-21.cdr
107
flame
© ISF 2002
Examples to Minimise Distortion
Figure 8.21
9. Welding Defects
9. Welding Defects
109
Figures 9.1 to 9.4 give a rough survey about the classification of welding defects to DIN 8524. This standard does not classify existing welding defects according to their origin but only to their appearance.
Figure 9.1
Figure 9.3
Figure 9.2
9. Welding Defects
110
A distinction of arising defects by their origin is shown in Figure 9.5. The development of the most important welding defects is explained in the following paragraphs.
Lack of fusion is defined as unfused area between weld metal and base material or previously welded layer. This happens when the base metal or the previous layer are not completely
or
insufficiently
molten. Figure 9.6 explains the influence of welding parameters on the development of lack of fusion. In Figure 9.4
the upper part, arc characteristic lines of MAG welding are shown using CO2 and mixed gas. The welding voltage depends on welding current and is selected according to the joint type. With present tension, the welding current is fixed by the wire feed
speed
(thus
also
melting rate) as shown in the middle part of the figFigure 9.5
ure.
Melting rate (resulting from selected welding parameters) and welding speed define the heat input. As it can be changed within certain limits, melting rate and welding speed do not limit each other, but a working range is created (lower part of the figure). If the heat input is too low, i.e. too high welding speed, a definite melting of flanks cannot be ensured. Due to the
9. Welding Defects
111
poor power, lack of fusion is the result. With too high heat input, i.e. too low welding speed, the weld pool gets too large and starts to flow away in the area in front of the arc. This effect prevents a melting of the base metal. The arc is not directed into the base metal, but onto the weld pool, and flanks are not entirely molten. Thus lack of fusion may occur in such areas.
Figure 9.6
Figure 9.7
Figure 9.7 shows the influence of torch position on the development of weak fusion. The upper part of the figure explains the terms neutral, positive and negative torch angle. Compared with a neutral position, the seam gets wider with a positive inclination together with a slight reduction of penetration depth. A negative inclination leads to narrower beads. The second part of the figure shows the torch orientation transverse to welding direction with multi-pass welding. To avoid weak fusion between layers, the torch orientation is of great importance, as it provides a reliable melting and a proper fusion of the layers. The third figure illustrates the influence of torch orientation during welding of a fillet weld. With a false torch orientation, the perpendicular flank is insufficiently molten, a lack of fusion occurs. When welding an I-groove in two layers, it must be ensured that the plate is com-
9. Welding Defects
112
pletely fused. A false torch orientation may lead to lack of fusion between the layers, as shown in the lower figure. Figure 9.8 shows the influence of the torch orientation during MSG welding of a rotating workpiece. As an example, the upper figure shows the desired torch orientation for usual welding speeds. This orientation depends on parameters
like
workpiece
diameter
and
thickness,
groove
shape,
melting
rate, and welding speed.
Figure 9.8
The lower figure illustrates variations of torch orientation on seam formation. A torch orientation should be chosen in such a way that a solidification of the melt pool takes place in 12 o'clock position, i.e. the weld pool does not flow in front or behind of the arc. Both may cause lack of fusion.
In contrast to faulty fusion, pores in the weld metal due to their globular shape are less critical, provided that their size does not exceed a certain value. Secondly, they must occur isolated and keep a minimum distance from each other. There are two possible mechanisms to develop cavities in the weld
Figure 9.9
metal: the mechanical and the metallurgical pore formation. Figure 9.9 lists causes of a mechanical pore formation as well as possibilities to avoid them. To over-weld a cavity (lack of
9. Welding Defects
113
fusion, gaps, overlaps etc.) of a previous layer can be regarded as a typical case of a mechanical pore formation. The welding heat during welding causes a strong expansion of the gasses contained in the cavity and consequently a development of a gas bubble in the liquid weld metal. If the solidification is carried out so fast that this gas bubble cannot raise to the surface of the weld pool, the pore will be caught in the weld metal.
a) X-ray photograph
b) Surface cross-section
c) Transverse section br-er09-10.cdr
© ISF 2002
Mechanical Pore Formation
Figure 9.10
Figure 9.11 Figure 9.10 shows a X-ray photograph
of
a
pore
which developed in this way, as well as a surface and a transverse section. This pore formation shows its typical pore position at the edge of the joint and at the fusion li ne of the top layer.
Figure 9.12
9. Welding Defects
114
Figure 9.11 summarises causes of and measures to avoid a metallurgical pore formation. Reason of this pore formation is the considerably increased solubility of the molten metal compared with the solid state. During solidification, the transition of liquid to solid condition causes a leapwise reduction of gas solubility of the steel. As a result, solved gasses are driven out of the crystal and are ena) X-rax photograph
riched as a gas bubble ahead of the solidification
front.
With
a
slow
growth
of
the
crystallisation front, the bubbles have enough time to raise to the surface of the weld pool, Figure 9.12 upper part. Pores will not be develb) surface section
oped. However, a higher solidification speed may lead to a case where gas bubbles are passed by the crystallisation front and are trapped as pores in the weld metal, lower part of the figure.
c) transverse section br-er09-13.cdr
© ISF 2002
Metallurgical Pore Formation
Figure 9.13 shows a X-ray photograph, a surface and a transverse section of a seam with
Figure 9.13 metallurgical pores. The evenly distributed pores across the seam and the accumulation of pores in the upper part of the seam (transverse
section)
are
typical.
Figure 9.14 shows the ways of ingress of gasses into the weld pool as an example
during
MAG
welding. A pore formation Figure 9.14
9. Welding Defects
115
is mainly caused by hydrogen and nitrogen. Oxygen is bonded in a harmless way when using universal electrodes which are alloyed with Si and Mn.
© ISF 2002
br-er09-15.cdr
Classification of Cracks to DIN 8524 Part 3 Figure 9.15
9. Welding Defects
116
Figure 9.15 classifies cracks to DIN 8524, part 3. In contrast to part 1 and 2 of this sta ndard, are cracks not only classified by their appearance, but also by their development.
Figure
9.16
allocates
cracks according to their appearance
during
the
welding heat cycle. Principally there is a distinction between the group 0010 (hot
cracks)
(cold cracks).
Figure 9.16 A model of remelting development and solidification cracks is shown in Figure 9.17. The upper part illustrates solidification conditions in a simple case of a binary system, under the provision that a complete concentration balance takes place in the melt ahead of the solidification front, but no diffusion takes place in the crystalline solid. When a melt of a composition C0 cools down, a crystalline solid is formed when the liquidus line is reached. Its concentration can be taken from the solidus line. In the course of the ongoing solidification, the rest of molten metal is enriched with alloy elements in accordance with the liquidus line. As defined in the beginning, no diffusion of alloy elements in the already solidified crystal takes place, thus the crystals are enriched with alloy elements much slower
Figure 9.17
and
0020
9. Welding Defects
117
than in a case of the binary system (lower line). As a result, the concentration of the melt exceeds the maximum equilibrium concentration (C 5), forming at the end of solidification a very much enriched crystalline solid, whose melting point is considerably lower when compared with the firstly developed crystalline solid. Such concentration differences between first and last solidified crystals are called segregations. This model of segregation development is very much simplified, but it is sufficient to understand the mechanism of hot crack formation. The middle part of the figure shows the formation of solidification cracks. Due to the segregation effects described above, the melt between the crystalline solids at the end of solidification has a considerably decreased solidus temperature. As indicated by the black areas, rests of liquid may be trapped by dendrites. If tensile stresses exist (shrinking stress of the welded joint), the liquid areas are not yet able to transfer forces and open up. The lower part of the figure shows the development of remelting cracks. If the base material to be welded contains already some segregations whose melting point is lower than that of the rest of the base metal, then these zones will melt during welding, and the rest of the material remains solid (black areas). If the joint is exposed to te nsile stress during solidification, then these areas open up (see above) and cracks occur. A hot cracking tendency of a steel is above all promoted by sulphur and phosphorus, because these elements form with iron very low melting phases (eutectic point Fe-S at 988°C) and these elements segregate i ntensely. In addition, hot crack te ndency increases with increasing melt interval.
As shown in Figure 9.18, also the geometry of the groove is important for hot crack tendency. With narrow, deep grooves a crystallisation takes place of all sides of the bead, entrapping the remaining melt in the bead centre. With the occurrence
of
shrinking
stresses, hot cracks may Figure 9.18
develop. In the case of flat beads as shown in the
9. Welding Defects
118
middle part of the figure, the
remaining
melt
solidifies at the surface of the bead. The melt cannot be trapped, hot cracking is not possible. The case in figure c shows no adva ntage, because a remelting crack may occur in the centre (segregation zone) of the first layer during welding the second layer.
The example of a hot
Figure 9.19
crack in the middle of a SA weld is shown in Figure 9.19. This crack developed due
to
the
unsuitable
groove geometry.
Figure 9.20 shows an example of a remelting crack which started to develop in a segregation zone of the base metal and spread up to the bead centre. The section shown in Figure 9.21 is similar to case c in Figure 9.18. One can
Figure 9.20
clearly see that an existing crack develops through the follo wing layers during over-welding. Figure 9.22 classifies cold cracks depending on their position in the weld metal area. Such a classification does not provide an explanation for the origin of the cracks.
9. Welding Defects
Figure 9.21
119
Figure 9.22
Figure 9.23 shows a summary of the three main causes of cold crack formation and their main influences. As explained in previous chapters, the resulting welding microstructure depends on both, the composition of base and filler materials and of the cooling speed of the joint. An unsatisfactory structure composition promotes very much the formation of cold cracks (hardening by martensite).
Another cause for increased cold crack susceptibility is a higher hydrogen content. The hydrogen content is very much influenced by the condition of the welding filler material (humidity of electrodes or flux, lubricating grease on welding wire etc.) and by humidity on the groove edges. The cooling speed is also important because it determines the remaining time for hydrogen effusion out of the bead, respectively how much hydrogen remains in the weld. A measure is t8/1 because only below 100°C a hydrogen e ffusion stops.
9. Welding Defects
Figure 9.23
120
Figure 9.24
A crack initiation is effected by stresses. Depending on material condition and the two already mentioned influencing factors, even residual stresses in the workpiece may actuate a crack. Or a crack occurs only when superimpose of residual stresses on outer stress.
Figure 9.24 shows typical cold cracks in a workpiece. An increased hydrogen content in the weld metal leads to an increased cold crack tendency. Mechanisms of hydrogen cracking were not completely understood until today. However, a spontaneous occurrence is typical of hydrogen cracking. Such cracks do not appear directly after welding but hours or even days after cooling. The weld metal hydrogen content depends on humidity of the electrode coating (manual metal arc welding) and of flux (submerged arc welding).
9. Welding Defects
121
Figure 9.25 shows that the moisture pick-up of an electrode coating greatly depends on ambient conditions and on the type of electrode. The upper picture shows that during storage of an electrode type the water content of the coating depends on air humidity. The water content of the coating of this electrode type advances to a maximum value with time. The lower picture shows that this behaviour does not
apply
to
all
electrode
types.
The
characteristics of 25 welding electrodes stored under identical conditions are plotted here. It can clearly be seen that a behaviour as shown in the upper picture applies only to some electrode types, but basically a very different behaviour in connection with storage Figure 9.25
can be noticed.
In practice, such constant storage conditions are not to be found, this is the reason why electrodes are backed before welding to limit the water content of the coating. Figure 9.26 shows the effects of this measure. The upper curve shows the water content of the coating of electrodes which were stored at constant air humidity before
Figure 9.26
rebaking. Humidity values after rebaking are plotted in the lower curve. It can be seen that even electrodes stored under
9. Welding Defects
122
very damp conditions can be rebaked to reach acceptable values of water content in the coating. Figure 9.27 shows the influence of cooling speed and also the preheat temperature on hydrogen content of the weld metal. The values of a high hygroscopic
cellulose-coated
electrode are considerably worse than of a basiccoated one, however both show the same tendency: increased cooling speed Figure 9.27
leads to a raise of diffusible hydrogen content in
weld metal. Reason is that hydrogen can still effuse all the way down to room temperature, but diffusion speed increases sharply with temperature. The longer the steel takes to cool, the more time is available for hydrogen to effuse out of the weld metal even in higher quantities. The table in Figure 9.28 shows an assessment of the quantity of diffusible hydrogen in weld metal according to DIN 8529.
Based on this assessment, a classification of weld metal to DIN 32522 into groups depending on hydrogen is carried out, Figure 9.29. Figure 9.28
9. Welding Defects
123 A cold crack development can
be
followed-up
by
means of sound emission measurement. Figure 9.30 represents the result of such a measurement of a welded
component.
A
solid-borne sound microphone is fixed to a component which measures the sound pulses generated by crack development. The
Figure 9.29
intensity of the pulses pro-
vides a qualitative assessment of the crack size. The observation is carried out without applying an external tension, i.e. cracks develop only caused by the internal residual stress condition. Figure 9.32 shows that most cracks occur relatively short after welding. At first this is due to the cooling process. Ho wever, after completed cooling a multitude of deve loping sounds can be registered. It is remarkable that the intensity of late occurring pulses is especially high. This behaviour is typical for hydrogen induced crack fo rmation.
Figure 9.31 shows a characteristic occurrence of lamellar cracks (also called lamellar tearing). This crack type occurs typically during stressing a plate across its thickness (perpendicular to rolling direction). The upper picture shows joint types which are very much at risk to formation of such cracks. The two lower
pictures
show
the
cause of that crack fo rmation. During steel production, a formation of segregation cannot be avoided due to Figure 9.30
9. Welding Defects
124
the casting process. With follo wing production steps, such segregations are stretched in the rolling direction. Zones enriched and depleted of alloy elements are now close together. These concentration diffe rences influence the transformation behaviour of the individual zones. During cooling, zones with enriched alloy elements develop a different microstructure than depleted zones. This effect which can be well recognised in Figure 9.31, is called structure banding. In practice, this formation can be hardly avoided. Banding in plates is the reason for worst mechanical properties perpendicular to rolling direction. This is caused by a different mechanical behaviour of different microstructures. When stressing lengthwise and transverse to rolling
direction,
the
individual structure
bands may support each other and a mean
br-er09-31.cdr
© ISF 2002
strength is provided. Such support cannot be obtained perpendicular to rolling direction, thus the strength of the
Figure 9.31 workpiece is that of the weaker microstructure areas. Consequently, a lamellar crack propagates through
weaker
micro-
structure areas, and partly a jump into the next band takes place.
Figure 9.32 illustrates why such t-joints are particularly vulnerable. DependFigure 9.32
ing on joint shape, these welds show to some extent
9. Welding Defects
125
a considerable shrinking. A welded construction which greatly impedes shrinking of this joint, may generate stresses perpendicular to the plane of magnitude above the tensile strength. This can cause lamellar tearing.
Precipitation cracks occur mainly during stress relief heat treatment of welded components. They occur in the coarse grain zone close to fusion line. As this type of cracks occurs often during post weld heat treatment of cladded materials, is it also called undercladding crack, Figure 9.33.
Especially susceptible are steels which contain alloy elements with a precipitation hardening effect (carbide developer like Ti, Nb, V). During welding such steels, carbides are dissolved in an area close to the fusion line. During the following cooling, the carbide developers are not completely re-precipitated.
If a component in such a condition is stress relief heat treated, a re-precipitation of carbides takes place (see hot ageing, chapter 8). With this re-precipitation, precipitation-free zones may develop along grain boundaries, which have a considerably lower deformation stress limit compared with strengthened areas. Plastic deformations during stress relieving are carried out almost only in these areas, causing the cracks shown in Figure 9.33.
Figure 9.33
10. Testing of Welded Joints
10. Testing of Welded Joints
Ls
127 The basic test for determination of material a
S S
S S
in test area
in test area
S
S
b
b1
S
S
Generally, it is carried out using a round
L0 Lc
r
behaviour is the tensile test. specimen. When determining the strength of a welded joint, also standardised flat speci-
Lt
total length head width
Lt b1
width of parallel length
plates
b
tubes
b
1 2
depends on test unit b + 12 12 with a £ 2 25 with a > 2 6 with D £ 50 12 with 50 < D £ 168,3 ³ L S + 60 ³ 25
Lc parallel length ) ) radius of throat r ) for pressure welding and beam welding, L S = 0. 2 ) for some other metallic materials (e.g.aluminium, copper and their alloys) __ L c ³ L S +100 may be required
mens are used. Figure 10.1 shows both standard specimen shapes for that test. A specimen is ruptured by a test machine while the actual force and the elongation of the
d1
d
S
S
1
r
is typical for this test, Figure 10.2.
L0 = measurement length (L0 = k ÖS0 with k = 5,65) Lt = total length S0 = initial cross-section within test length
br-er10-01.cdr
ment values, tension σ and strain ε are calculated. If σ is plotted over ε, the drawn diagram
LO LC Lt d = specimen diameter d1 = head diameter depending on clamping device LC = test length = L0 + d/2 r = 2 mm
specimen is measured. With these measure-
Normally, if a steel with a bcc lattice structure © ISF 2002
Flat and Round Tensile Test Specimen to EN 895, EN 876, and EN 10 002
is tested, a curve with a clear yield point is obtained (upper picture). Steels with a fcc lattice structure show a curve without yield
Figure 10.1
point. The most important characteristic values
s
which are determined by this test are: yield stress ReL, tensile strength Rm, and elongation
Rm ReH Rel sf
A. To determine the deformability of a weld, a e
ALud
bending test to DIN EN 910 is used, Figure
Ag A
10.3. In this test, the specimen is put onto two
s Rm
supporting rollers and a former is pressed
RP0,2 RP0,01 sf
through between the rollers. The distance of the supporting rollers is Lf = d + 3a (former diameter + three times specimen thickness). e
0,2 % 0,01 % Ag
is observed. If a surface crack develops, the
A br-er10-02.cdr
© ISF 2002
Stress-Strain Diagram With and Without Distinct Yield Point
Figure 10.2
The backside of the specimen (tension side) test will be stopped and the angle to which the specimen could be bent is measured. The
10. Testing of Welded Joints
128
test result is the bending angle and the diameter of the used former. A bending angle of 180° is reached, if the specimen is pressed through the supporting rollers without development of a crack. In Figure 10.3 specimen shapes of this test are shown. Depending on the direction the weld is bent, one distinguishes (from top to bottom) transverse, side, and longitudinal bending specimen. The tension side of all three speci-
section A-B tension side
A
r
former
b
men types is machined to
supporting roller
r a
the test through notch ef-
B Lt
a
d
eliminate any influences on
bending specimen
section A-B Lf l Lt
r
b
fects. Specimen thickness of
A
r
tension side
thickness.
Side
plate
b
the
bending
r
is
tension side
r
specimens
B Lt
a
transverse and longitudinal
a
specimens are normally only
l Lt d D a r b
D
distance of supporting rollers specimen length former diameter supporting roller diameter: 50 mm specimen thickness radius of specimen edge specimen width
Lt
br-er10-03.cdr
used with very thick plates,
Bending Specimens to EN 910
here the specimen thickness Figure 10.3
is fixed at 10 mm.
A determination of the toughness of a material or welded joint is carried out with the notched bar impact test. A cuboid specimen with a V-notch is placed on a support and then hit by a pendulum ram of the im55
10
pact testing machine (with
8
10
r = 0,25
very tough materials, the
45 J Charpy impact energy
40
specimen will be bent and
45°
40
average values maximaum values minimum values
35
D im e nsio ns leng th width hight notc h angle
No m inal s ize 55 mm 10 mm 10 mm 45°
± ± ± ±
To leranc e 0,6 0 mm 0,1 1 mm 0,0 6 mm 2°
thic knes s in notch g roove notc h rad ius notc h d is tanc e from end of s p ecim en angle b etwee n no tch axis and long itudinal axis
8 0 ,2 5
mm mm
± 0,0 6 ± 0,0 2 5
mm mm
2 7,5
mm
± 0,4 2
mm
30 25 20 15
90°
± 2°
drawn through the supports). The used energy is measured.
Figure
10.4
represents sample shape, notch
shape
(Iso-V-
10 -80
-60
br-er10-04.cdr
-40 -20 0 Temperature
20 °C 40
Charpy Impact Test Specimen and Schematic Representation of Test Results
Figure 10.4
specimen), and a schematic presentation of test results.
10. Testing of Welded Joints
129 Three specimens are tested at each test tem-
b
Designation
VWS a/b
Dicke
a
RL
VWS a/b (fusion weld)
Fusion line/bonding zone
perature, and the average values as well as b
Weld centre
Designation
RL
the range of scatter are entered on the impact
a
Dicke
b
b
energy-temperature diagram (AV-T curve). VWT 0/b
VHT 0/b
This graph is divided into an area of high im-
a
b
b
pact energy values, a transition range, and an VHT a/b
VWT a/b
a
area of low values. A transition temperature is
VWT 0/b
b
b
a
VHT a/b
b
b
VWT a/b
drop of toughness values. When the tempera-
a RL
RL
assigned to the transition range, i.e. the rapid ture falls below this transition temperature, a
VHT a/b
transition of tough to brittle fracture behaviour
a RL
a RL
V = Charpy-V notch W = notch in weld metal; reference line is centre line of weld H = notch in heat affected zone; reference line is fusion line or bonding zone (notch should be in heat affected zone) S = notched area parallel to surface T = notch through thickness a = distance of notch centre from reference line (if a is on centre line of weld, a = 0 and should be marked) b = distance between top side of welded joint and nearest surface of the specimen (if b is on the weld surface, then b = 0 and should be marked) br-er10-05.cdr
takes place. As this steep drop mostly extends across a certain area, a precise assignment of transi© ISF 2002
Position of Charpy-V Impact Test Specimen in Welded Joints to EN 875
tion temperature cannot be carried out. Following DIN 50 115, three definitions of the transition temperature are useful, i.e. to fix TÜ
Figure 10.5
to:
1.) a temperature where the level of impact values is half of the level of the high range, 2.) a temperature, where the fracture area of the specimen shows still 50% of tough fracture behaviour 3.) a temperature with an impact energy value of 27 J. Figure 10.5 illustrates a specimen position and notch position related to the weld according to DIN EN 875. By modifying the notch position, the impact energy of the individual areas like HAZ, fusion line, weld metal, and base metal can be determined in a relatively accurate way. Figure 10.6 presents the influence of various alloy elements on the AV-T - curve. Three basically different influences can be seen. Increasing manganese contents increase the impact values in the area of the high level and move the transition temperature to lower values. The values of the low levels remain unchanged, thus the steepness of the drop becomes clearer with increasing Mn-content. Carbon acts exactly in the opposite way. An increasing carbon content increases the transition temperature and lowers the values of the high level, the steel becomes more brittle. Nickel decreases slightly the values of the high level, but increases the
10. Testing of Welded Joints
130 values of the low level with increasing con-
specimen position: core longitudinal
J
tent. Starting with a certain Nickel content
specimen shape: ISO V
(depends also from other alloy elements), a
300 2% Mn
steep drop does not happen, even at lowest
1% Mn
200
0,5% Mn
temperature the steel shows a tough fracture
Charpy impact energy AV
100
behaviour. 0% Mn
27 200
In Figure 10.7, the AV-T – curves of some
J 100
27
13% Ni 8,5% 5% 3,5%
2% Ni
commonly used steels are collected. These
0% Ni
curves are marked with points for impact en-
200
ergy values of AV = 27 J as well as with points
0,1% C
J
where the level of impact energy has fallen to
100 0,4% C
half of the high level. It can clearly be seen
0,8% C
27 -150
-100
-50 0 Temperature
50
that mild steels have the lowest impact en-
°C 100 © ISF 2002
br-er10-06.cdr
Influence of Mn, Ni, and C on the Av-T-Curve
ergy values together with the highest transition temperature. The development of finegrain structural steels resulted in a clear im-
Figure 10.6
provement of impact energy values and in
addition, the application of such steels could be extended to a considerably lower temperature range. With the example of the steels St E 355 and St E 690 it is clearly visible that an increase of strength goes mostly hand in hand with a decrease of the impact energy level. Another improvement showed the application of a thermomechanical treatment (controlled rolling during heat treatment). The applispecimen position: weld centre, notch parallel to surface specimen shape: standard specimen with V-notch J
sulted in an increase of
300
strength and impact energy values together with a parallel saving of alloy elements. To make a comparison, the AV-T - curve of the cryogenic
Charpy impact energy AV
cation of this treatment re-
X8Ni9 S460M
S355N
S690N 200 S235J2G3
S355J2G3
100
27
and high alloyed steel X8Ni9
-150
-100
-50 Temperature
0
50
was plotted onto the diabr-er10-07.cdr
gram. The material is tested AV-T Curves of Various Steel Alloys
Figure 10.7
°C
100
10. Testing of Welded Joints
131
under very high test speed in the impact enP
C
1,2h ± 0,25
about crack growth and fracture mechanisms.
0,55h ± 0,25
C
ergy test, thus there are no reliable findings
P a
Figure 10.8 shows two commonly used
b
CT - specimen
L h 1,25h ± 0,13
specimen shapes for a fracture mechanics
specimen height h = 2b ± 0,25 specimen width b total crack length a = (0,50 ± 0,05)h test load P
test to determine crack initiation and crack
a
h
growth. The lower figure to the right shows a possibility how to observe a crack propaga-
2,1h
2,1h
b
S
tion in a compact tensile specimen. During the test, a current I flows through the speci-
SENB -specimen 3PB
specimen width b
bearing distance S = 4h
sample height h = 2b ± 0,05
total crack length a = (0,50 ± 0,05)h
F,U
crack initiation
U F
men, and the tension drop above the notch is
UE,aE U
measured.
UO V
As soon as a crack propagates through the
V
br-er10-08.cdr
© ISF 2002
Fracture Mechanics Test Sample Shape and Evaluation
material, the current conveying cross section decreases, resulting in an increased voltage Figure 10.8
drop. Below to the left a measurement graph
of such a test is shown. If the force F is plotted across the widening V, the drawn curve does not indicate precisely the crack initiation. Analogous to the stress-strain diagram, a decrease of force is caused by a reduction of the stressed cross-section. If the voltage drop is plotted over the force, then the start of crack initiation can be determined with suitable accuracy, and the crack propagation can F
be observed.
F
Another typical characteristic of material behaviour is h
the hardness of the workpiece. Figure 10.9 shows hardness test methods to
d
Brinell
(standardised
to
d
d1
2
DIN 50 351) and Vickers to Brinell, a steel ball is
br-er-10-09.cdr
Hardness Testing to Brinell and Vickers
Figure 10.9
(DIN 50 133). When testing pressed with a known load
10. Testing of Welded Joints
132
to the surface of the tested workpiece. The diameter of the resulting impression is measured and is a magnitude of hardness. The hardness value is calculated from test load, ball diameter, and diameter of rim of the impression (you find the formulas in the standards). The hardness information contains in addition to the hardness magnitude the ball diameter in mm, applied load in kp and time of influence of the test load in s. This information is not required for a ball diameter of 10 mm, a test load of 3000 kp (29420 N), and a time of influence of 10 to 15 s. This hardness test method may be 3 6
2
7
10
3
6
7
0
7
8,9 10
3 10
specimen surface
6
130
30 0
hardness scale
6 hardness scale
100
reference level for measurement
4 5 3 8
130 30 0
specimen surface
0,200 mm
Instead of a ball, a diamond pyramid is
1
3
100 0
Hardness testing to Vickers is analogous. This method is standardised to DIN 50133.
4 5 3 8
0,200 mm
(Brinell Hardness Number).
0,200 mm
1
0,200 mm
used only on soft materials up to 450 BHN
8,9
reference level for measurement
7 10
pressed into the workpiece. The lengths of the two diagonals of the impression are
Terms
Abbreviation
ball diameter = 1,5875 mm ( 1/16 inch)
-
cone angle = 120°
2
-
radius of curvature of cone tip = 0,200 mm
3
F0
test preload
4
F1
test load
5
F
total test load = F0 + F1
6
t0
penetration depth in mm under test preload F0. This defines the reference level for measurement of tb.
The impressions of the test body are always
7
t1
total penetrationn depth in mm under test load F1
8
tb
resulting penetration depth in mm, measured after release of F1 to F0
geometrically similar, so that the hardness
9
e
resulting penetration depth, expressed in units of 0,002 mm: tb / 0,002
10
HRC HRA
measured and the hardness value is calculated from their average and the test load.
1
value is normally independent from the size of the test load. In practice, there is a hard-
Rockwell hardness = 100 - e
HRB HRF
e =
Rockwell hardness = 130 - e
br-er10-10.cdr
© ISF 2002
Hardness Test to Rockwell
ness increase under a lower test load because of an increase of the elastic part of the deformation.
Figure 10.10
Hardness testing to Vickers is almost universally applicable. It covers the entire range of materials (from 3 VHN for lead up to 1500 VHN for hard metal). In addition, a hardness test can be carried out in the micro-range or with thin layers. Figure 10.10 illustrates a hardness test to Rockwell. In DIN 50103 are various methods standardised which are based on the same principle. With this method, the penetration depth of a penetrator is measured.
10. Testing of Welded Joints
133
At first, the penetrator is put on the workpiece by application of a pre-test load. The purpose is to get a firm contact between workpiece and penetrator and to compensate for possible play of the device. Then the test load is applied in a shock-free way (at least four times the pre-force) and held for a certain time. Afterwards it is released to reach minor load. The remaining penetration depth is characteristic for the hardness. If the display instrument is suitably scaled, the hardness value can be read-out directly. All hardness test methods to Rockwell use a ball (diameter 1.5875 mm, equiv. to 1/16 Inch) or a diamond sphero-conical penetrator (cone angle 120°) as the penetrating body. There are differences in size of pre- and test load, so different test methods are scaled for different hardness ranges. The most commonly used scale methods are Rockwell B and C. The most considerable advantage of these test methods compared with Vickers and Brinell are the low time duration and a possible fully-automatic measurement value recognition. The disadvantage is the reduced accuracy in contrast to the other methods. Measured hardness numbers are only comparable under identical conditions and with the same test method. A comparison of hardness values which were determined with different methods can only be carried out for similar materials. A conversion of hardness values of different methods can be carried out piston
for steel and cast steel according to a table in DIN 50150. A relation of hardness and tensile strength is also given in that table. All the hardness test methods described above require a coupon which must be taken from the
reference bar
workpiece and whose hardness is then determined in a test machine. If a workpiece on-site is to be tested, a dynamical hardness test
specimen
method will be applied. The advantage of these methods is that measurements can be taken
br-er10-11.cdr
on completed constructions with handheld
© ISF 2002
Poldi - Hammer
units in any position. Figure 10.11 illustrates a Figure 10.11
10. Testing of Welded Joints
134
hardness test using a Poldi-Hammer. With this (out of date) method, the measurement is carried out by a comparison of the workpiece hardness with a calibration piece. For this purpose a calibration bar of exactly determined hardness is inserted into the unit, which is held by a spring force play-free between a piston and a penetrator (steel ball, 10 mm diameter). The unit is put on the workpiece to be tested. By a hammerblow to the piston, the penetrator penetrates the workpiece and the calibration pin simultaneously. The size of both impressions is measured and with the known hardness of the calibration bar the hardness of the workpiece can be determined. However, there are many sources of errors with this method which may influence the test result, e.g. an inclined resting of the unit on the surface or a hammerblow which is not in line with the device axis. The major source of errors is the measurement of the ball impression on the workpiece. On one hand, the edge of the impression is often unsharp because of the great ball diameter, on the other hand the measurement of the impression using magnifying glasses is subjected to serious errors. Figure 10.12 shows a modern measurement method which works with ultrasound and combines a high flexibility with easy handling and high accuracy. Here a test tip is pressed manually against a workpiece. If a defined test load is passed, a spring mechanism inside the test tip is triggered and the measurement starts. Test force
The measurement principle is based on a measurement of damping characteristics in 5 kp
5.0
the steel. The measurement tip is excited to
kp
emit ultrasonic oscillations by a piezoelectric
4.0
crystal. The test tip (diamond pyramid) pene3.0
trates the workpiece under the test pressure 2.0
caused by the spring force. With increasing Federweg
penetration depth the damping of the ultrasonic oscillation changes and consequently the frequency. This change is measured by the device. The damping of the ultrasonic os- little work on surface preparation of specimens (test force 5 kp) - Data Logger for storage of several thousands of measurement points - interfaces for connection of computers or printers - for hardness testing on site in confined locations
br-er10-12.cdr
© ISF 2002
cillation depends directly on penetration depth thus being a measure for material hardness. The display can be calibrated for all commonly used measurement methods, a measurement is carried out quickly and easily.
Figure 10.12
10. Testing of Welded Joints
135
Measurements can also be carried out in confined spaces. This measurement method is not
pulsation range (compression)
Application
Dye penetrant method
σm = σa
σm > σ a
crack is free, surface is clean
σm < σa
compression -
+ tension
Description σm = 0
σ m < σa
σm = σa
σ m > σa
yet standardised.
time
crack and surface with penetrant liquid cleaned surface, dye penetrant liquid in crack
pulsation range (tension)
alternating range
all materials with surface cracks
surface with developer shows the crack by coloring
Wöhler line Magnetic particle testing
II
A workpiece is placed between the poles of a magnet or solenoid. Defective parts disturb the power flux. Iron particles are collected.
I
III
σD
Stress σ
failure line
Surface cracks and cracks up to 4 mm below surface. However: Only magnetizable materials and only for cracks perpendicular to power lines
0 1 10 102 103 104 105 106 Fatigue strength (endurance) number lg N
107
I area of overload with material damage II area of overload without material damage III area of load below fatigue strength limit
br-er10-13.cdr
© ISF 2002
br-er10-14.cdr
© ISF 2002
Fatigue Strength Testing
Figure 10.13
Figure 10.14
To test a workpiece under oscillating stress, the fatigue test is standardised in DIN 50100. Mostly a fatigue strength is determined by the Wöhler procedure. Here some specimens (normally 6 to 10) are exposed to an oscillating stress and the number of endured oscillations until rupture is determined (endurance number, number of cycles to failure). Depending on where the specimen is to be stressed in the range of pulsating tensile stresses, alternating stresses, or pulsating compressive stresses, the mean stress (or sub stress) of a specimen group is kept constant and the stress amplitude (or upper stress) is varied from specimen to specimen, Figure 10.13. In this way, the stress amplitude can be determined with a given medium stress (prestress) which can persist for infinite time without damage (in the test: 107 times). Test results are presented in fatigue strength diagrams (see also DIN 50 100). As an example the extended Wöhler diagram is shown in Figure 10.13. The upper line, the Wöhler line, indicates after how many cycles the specimen ruptures under tension amplitude σa. The
10. Testing of Welded Joints
Description
136
Application
X-ray or isotope radiation penetrate a workpiece. The thicker the workpiece, the weaker the radiation reaching the underside.
W ire diameter
Mainly for defects with orientation in radiation direction.
Tolerated deviation
mm 3,2 2,5 2 1,6 1,25 1 0,8 0,63 0,5 0,4 0,32 0,25 0,2 0,16 0,125 0,1
¬
−
W ire number
mm 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16
± 0,03
± 0,02
± 0,01
± 0,005
° Abbreviation
®
FE 1/7
1 to 7
FE 6/12 FE 10/16 CU 1/7
−
CU 10/16
¯
AL 1/7 AL 6/12
workpiece
AL 10/16
® film (displayed in distance from workpiece) ¯ defect in radiation direction; difficult to identify (flank lack of fusion) ° defect in radiation direction; easy to identify br-er10-15.cdr
© ISF 2002
W ire length mm 50
6 to 12
50 or 25
10 to 16
50 or 25
1 to 7
CU 6/12
¬ radiation source
W ire number to Table 1
6 to 12 10 to 16
Material groups to be tested
mild steel
iron materials
copper
copper, zink, tin and its alloys
aluminium
aluminium and its alloys
50 50 50 or 25
1 to 7
50
6 to 12
50
10 to 16
W ire material
50 or 25
br-er10-16.cdr
© ISF 2002
Determination of Picture Quality Number to DIN 54105
Non-Destructive Test Methods Radiographic Testing
Figure 10.16
Figure 10.15
damage line indicates analogously, when a Description US-head generates high-frequency sound waves, which are transferred via oil coupling to the workpiece. Sound waves are reflected on interfaces (echo).
Application Mainly for defects with an orientation transverse to sound input direction.
damage to the material starts in form of cracks. Below this line, a material damage does not occur.
Ã
Test
À
methods
described
above
require
specimens taken out of the workpiece and a
Á
partly very accurate sample preparation. A testing of completed welded constructions is
Â
impossible, because this would require a deÄ À sound head Á oil coupling  workpiece à defect Ä ultrasonic test device Å radiation pulse Æ defect echo ³ backwall echo
Å
Æ
³
br-er10-17.cdr
© ISF 2002
Non-Destructive Test Methods Ultrasonic Testing II
Figure 10.17
struction of the workpiece. This is the reason why various non-destructive test methods were developed, which are not used to determine technological properties but test the workpiece for defects. Figure 10.14 shows
10. Testing of Welded Joints
137
two methods to test a workpiece for surface defects. Figure 10.15 illustrates the principle of radiographic testing which allows to identify also defects in the middle of a weld. The size of the minimum detectable defects depends greatly on the intensity of radiation, which must be adapted to the thickness of the workpiece to be radiated. As the film with documented defects does not permit an estimation of the plate thickness, a scale bar must be shown for estimation of the defect size. For that purpose, a plastic template is put on the workpiece before radiation which contains metal wires with different thickness and incorporated metallic marks, Figure 10.16. The size of the thinnest recognisable wire indicates the Figure 10.18
size of the smallest visible defect. Radiation testing provides information about the defect
position in the plate plane, but not about the position within the thickness depth. A clear advantage is the good documentation ability of defects. An information about the depth of the defect is provided by testing the workpiece with ultrasound. The principle is shown in Figures
10.17
and
10.18
(principle of a sonar). The br-er10-19.cdr
of
original
pulse, backwall and defect Ultrasonic Testing of Fillet Welds
Figure 10.19
display
echo is carried out with an oscilloscope.
10. Testing of Welded Joints
138
This method provides not only a perpendicular sound test, but also inaccessible regions can be tested with the use of so called angle testing heads, Figure 10.19.
Pores between 10 and 20 mm depth provide an unbroken echo sequence across the entire display starting from 10mm. The backwall echo sequence of 30 mm is not yet visible.
30
Wall thickness is below 40 mm. The roughness provides smaller and wider echos.
Echo sequence of 20 mm depth. The backwall is completely screened.
The perpendicular crack penetrating the material does not provide a display because the reflecting surface (tip of crack) is too small.
40 The oblique and rough defect from 20 to 30 mm provides a wide echo of 20 to 30 mm. Starting with SKW 4, an unbroken echo sequence follows. The inclination of the reflector is recornised by a change of the 1st echo when shifting the test head.
The oblique backwall reflects the soundwaves against the crack. this is the reason why an ‘impossible’ depth of 65 mm is displayed.
Echo sequence of 10 mm depth. The reflector in 30 mm depth is completely screened.
br-er10-20.cdr
© ISF 2002
br-er10-21.cdr
Defect Identification with Ultrasound
© ISF 2002
Defect Identification With Ultrasound
Figure .10.20
Figure 10.21 Figures 10.20 and 10.21 show macro section
schematically
the
display of various defects on an oscilloscope. A cor-
base material
50 µ
ferrite + perlite
coarse grain zone
bainite
rect interpretation of all the signals requires great experience,
2,5 mm
fine grain zone
ferrite + perlite
fusion line Steel: S355N (T StE 355) weld metal
bainite
because
the
shape of the displayed signals is often not so clear.
cast structure
br-er10-22.cdr
Metallographic Examination of a Weld
Figure 10.22 illustrates the potential of metallographic
Figure 10.22
examination. Grinding and
10. Testing of Welded Joints
139 etching with an acid makes the microstructure visible. The reason is that depending on structure and orientation, the individual grains react very differently to the acid attack thus 100
25
Fe
% Fe
% Cr
macrosection, i.e. without magnification, gives
Cr 20
60 40
15
20
10 % Ni
reflecting the light in a different way. The
80
a complete survey about the weld and fusion line, size of the HAZ, and sequence of solidification. Under adequate magnification, these
0 10
areas can still not be distinguished precisely,
8
Ni
however, an assessment of the developed
6 4
5
microstructure is possible.
2 0
0 200
mm
100
0
An assessment of the distribution of alloy
100
Distance from fusion line br-er10-23.cdr
elements across the welded joint can be car-
© ISF 2002
Micro-Analysis of the Transition Zone Base Material - Strip Cladding
ried out by the electron beam micro-analysis. An example of such an analysis is shown in
Figure 10.23
Figure 10.23. If a solid body is exposed to a
focused electron beam of high energy, its atoms are excited to radiate X-rays. There is a simple relation between the wave length of this radiation and the atomic number of the chemical elements. As the intensity of the radiation depends on the concentration of the elements, the chemical composition of the solid body can be concluded from a survey of the emitted
X-ray
qualitatively
and
spectrum quantita-
tively. A detection limit is
50
50
50
20 20
1. weld
about 0.01 mass % with this
50
20 20
2. weld 0 10
method. Microstructure areas of a minimum diameter
weld
of about 5 µm can be ana-
axis of bending former
weld
Agents: - electrolytic copper in the form of chips (min. 50 g/l test solution) - 100 ml H2SO4 diluted with 1 l water and then . 110 g CuSO 5 H2O are added
lysed. If the electron beam is
Test: The specimens remain for 15 h in the boiling test solution. Then the specimens are bent across a former up to an angle of 90° and finally examined for grain failure under a 6 to 10 times magnification.
moved across the specimen (or the specimen under the br-er-10-24.cdr
beam), the element distribu-
Strauß - Test
tion along a line across the Figure 10.24
axis of bending former
10. Testing of Welded Joints
140
solid body can be determined. Figure 10.23 presents the distribution of Ni, Cr, and Fe in the transition zone of an austenitic plating in a ferritic base metal. The upper part shows the related microsection which belongs to the analysed part. This microanalysis was carried out along a straight line between two impressions of a Vickers hardness test. The impressions are also used as a mark to identify precisely the area to be analysed. The so called Strauß test is 12
standardised in DIN 50 914. it serves to determine
80
web
the resistance of a weld
measurement points
tack welds
against intergranular corro-
base plate weld1
40
40
20
sion. Figure 10.24 shows the specimen shape which
a
a a
20
aa
a
a
12
weld2
is normally used for that
120
80
aa
test. In addition, some debr-er-10-25.cdr
Test of Crack Susceptibility of Welding Filler Materials to DIN 50129
tails of the test method are explained.
Figure 10.25 Figure 10.25 presents a specimen shape for testing the crack susceptibility of welding consumables. For this test, weld number 1 is welded first. The 2. weld is welded not later than 20 s in reversed direction after completion of the first weld. Throat thickness of weld 2 must be 20% below of weld 1. After cooling down, the beads are examined for cracks. If tensioning bolt hexagon nut min. M12 DIN 934
guidance plates
a tensioning plate specimen base body
cracks are found in weld 1, the test is void. If weld 1 is free from cracks, weld 2 is examined for crack with magnifying glasses. Then weld 1 is machined off and weld 2 is cracked by bend-
br-er-10-26.cdr
Tensioning Specimen for Crack Susceptibility Test
Figure 10.26
ing the weld from the root. Test results record any
10. Testing of Welded Joints
141
surface and root cracks together with information about position, orientation, number, and length. The welding consumable is regarded as 'non-crack-susceptible' if the welds of this test are free from cracks. Figure 10.26 presents two proposals for self-stressing specimens for plate tests regarding their hot crack tendency. Such tests are not yet standardised to DIN.
thermo couple electrode
cross-section
groove shape 60°
60°
welding direction
weld metal support plate
Wd./2 H
Wd.
2
implant
Hc
Wd./2
2 load temperature in °C
specimen shape
load in N
Tmax start
end crater
150
crack coefficient
C=
c
x 100 (in %)
800 500
1
2 3
4 5
sections 60 anchor weld
80 test weld
150 100 60 anchor weld
br-er10-27.cdr
t8/5
© ISF 2002
rupture time
br-er10-28.cdr
Tekken Test
Figure 10.27
time in s
© ISF 2002
Implant Test
Figure 10.28
There are various tests to examine a cold crack tendency of welded joints. The most important ones are the self-stressing Tekken test and the Implant test where the stress comes from an external source. In the Tekken test which is standardised in Japan, two plates are coupled with anchor joints at the ends as a step in joint preparation see Figure 10.27. Then a test bead is welded along the centre line. After storing the specimen for 48 hours, it is examined for surface cracks. For a more precise examination, various transverse sections are planned. The value to be determined is the minimum working temperature at which cracks no longer occur. The specimen shape simulates the conditions during welding of a root pass.
10. Testing of Welded Joints
142
The most commonly used cold crack test is the Implant test, Figure 10.28. A cylindrical body (Implant) is inserted into the bore hole of a support plate and fixed by a surface bead. After the bead has cooled down to 150°C the implant is exposed to a constant load. The time is measured until a rupture or a crack occurs (depending on test criterion 'rupture' or 'crack'). Varying the load provides the possibility to determine the stress which can be born for 16 hours without appearance of a crack or rupture. If a stress is specified to be of the size of the yield point as a requirement, a preheat temperature can be determined by varying the working temperature to the point at which cracks no longer appear. As explained in chapter 'cold cracks' the hydrogen content plays an important role for cold crack development. Figure 10.29 shows results of trials where the cold crack behaviour was examined using the Tekken and Implant test. Variables of these tests were hydrogen content of the weld metal and preheat temperature. The variation of the hydrogen content of the weld metal was carried out by different exposure to humidity (or rebaking) of the used stick electrodes. Based on the hydrogen content, the preheat temperature was increased test by test. Consequently, the curves of Figure 10.29 represent the limit curves for the related test. Specimens above these heat input: 12 kJ/cm basic coated stick electrode plate and support plate thickness: 38 mm
°C
cracks, below these curves
°C Implant-Test
150
Tekken-Test
100
50
cracks are present. Evi-
150
Rcr = Rp0,2 = 358 N/mm² Preheat temperature
Preheat temperature
curves remain free from
fractured starting cracks crack-free
20
dent for both graphs is that with
100
temperature 50
starting cracks crack-free
20 0
10
20
30
ml/ 40 100 g
increased
0
10
Diffusible hydrogen content br-er-10-29.cdr
Test Result Comparison of Implant and Tekken Test
20
30
ml/ 40 100 g
preheat
considerably
higher hydrogen contents are tolerated without any crack
development
be-
cause of the much better hydrogen effusion.
Figure 10.29 If both graphs are compared it becomes obvious that the tests produce slightly different findings, i.e. with identical hydrogen content, the determined preheat temperatures required for the avoidance of cracking, differ by about 20°C.
10. Testing of Welded Joints
143
Figure 10.30 illustrates a method to measure the diffusible hydrogen content in welds which is standardised in DIN 8572. Figure a) shows the burette filled with mercury before a specimen is inserted. The coupons are inserted into the opened burette and drawn with a magnet through the mercury to the capillary side (density of steel is lower than that of mercury, coupons surface). Then the burette is closed and evacuated. The hydrogen, which effuses of the coupons but does not diffuse through the mercury, collects in the capillary. The samples remain in the evacuated burette 72 hours for degassing. To determine the hydrogen volume the burette is ventilated and the coupons are removed from the capillary side. The volume of the effused hydrogen can be read out from the capillary; the height difference of the two mercury menisci, the air pressure, and the temperature provide the data to calculate the
norm
volume
to pump hydrogen under reduced pressure
under
VT
air pressure B
evacuated
standard
conditions.
This
capillary side
volume and the coupons
M
meniskus1
weight are used to calculate,
meniskus2 mercury
coupons
as measured value, the hydrogen volume in ml/100 g weld metal. This is the most
a) starting condition
b) during degassing
c) ventilated after degassing
br-er-10-30.cdr
commonly used method to determine
the
Burettes for Determination of Diffusible Hydrogen Content
hydrogen
content in welded joints.
Figure 10.30
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