Establishing Squeeze Time
[from PWL#102 - Section 3]
This is not the first time I pick up useful hints from the always interesting Q&A notes published by The Welding Journal (January 2012, page 20). I always recommend these informative pages.
A reader had asked where to find squeeze time data to spot weld cold rolled and galvanized steel, as it is not published along other schedule data in the RWMA Resistance Welding Manual. The Author, Roger Hirsch, suggested in his answer to find it by trial and error.
One should remind that squeeze time is the time interval between timer initiation and first application of current. It is designed to allow electrode movement and full force action deployment before performing the spot weld.
It appears that this interval is dependent on design details of each machine: because of this reason it cannot be a fixed datum to be published on tables.
Too short a squeeze time will manifest itself by the appearance of sparks from under the electrode. By simply increasing this time in parallel tests until sparking (almost) stops one can find the correct value (in cycles) needed for correct operation.
Note that any further increment does not add any advantages and wastes production time.
The alternative answer that involves calculations seems less attractive.
If the equipment has a differential pressure transducer that monitors the air pressure on both sides of the cylinder, one can establish the time when full pressure is applied to the electrodes. If a signal can then be used to start the current cycles, squeeze time is redundant and can be bypassed (set to zero).
Interested readers are referred to the original article indicated above.
How many weld repairs are allowed?
[From PWL#103 - Section 3]
You may recall having seen requirements limiting the number of allowed repairs performed on the same location, for the purpose of protecting the construction from weak spots likely to develop due to repeated weld repairs.
If you are bound by such a customer requirement you can either argue on its necessity or abide to it. But now you can at least find a reference to a research program intended to test if there is a reason supporting this request.
A group of researchers, knowing the limitations but unable to find on the subject precise indication of accepted Standards, undertake a testing program intended to verify, one way or the other, the influence of actual simulated cut and repair cycles on the properties of a welded joint.
You will find the summary in the February 2012 Issue of the Welding Journal at page 25.
In the specific case the test was performed on low carbon steel, ASTM A 283 GrB, flat plates 3/8 in. (9.5 mm) thick.
Test pieces were welded together using direct current, with the gas metal arc welding (GMAW) process, in the flat (1G) position, with a wire of 1.2 mm diameter, type ER 70S-6 per AWS Specification A5.18, recommended for welding low-carbon steel. The shielding gas was 75% argon and 25% carbon dioxide.
Six equal test pieces 200 mm wide ׳ 440 mm long were welded after having been prepared with a 60 deg. bevel, manually cut with an oxyacetylene torch and cleaned with a grinding disk. After the face of the specimens was welded, the root was backgauged with a file and rewelded.
The first specimen was put aside for testing. The remaining specimens were cut with the same methods and rewelded. The second specimen having one weld and one repair weld after cutting was put aside for testing. The other specimens followed the same procedure. The third had one weld and two repairs, the fourth one weld and three repairs, the fifth one weld and four repairs and the last one had one weld and five repairs.
Test pieces removed from the six specimens were tested for bending, ultimate tensile, impact, elongation, average grain size, and metallographic structure of the HAZ.
The results showed a remarkable homogeneity, without dramatic conditions likely to alert the researchers of dangerous worsening of basic properties. The conclusion reached in this program was that welding can be performed safely on the same area at least six times (one weld and five repairs) on low-carbon steel.
My comment is that the conclusions may be correct for applications reflecting in all details the procedures of this research. It would be risky to extend them to different situations.
Interested readers are urged to seek the original article reported above.
[From PWL#104 - Section 3]
From time to time I get queries trying to locate parameter changes capable of reducing distortion while welding. The following, besides other queries, illustrate the kind of questions on this argument:
1) - "I need to do some extensive aluminum welding on a cast outboard block for modification purposes. I have a Lincoln square wave tig 255. I'm experienced in welding from thin Alu
sheets to the much thicker ones and different parts etc.
I've been told to watch for warping/distortion when welding Aluminum that extensively.
My question now is what can I do to reduce these risks of distortion or warping?"
2) - "We have a storage tank (18000m3, floating roof, type of metal A283C, diameter of tank 44m approximately, height 13m). Thickness of bottom 9mm. Thickness of floating roof 7mm.
Thickness of shell (18,16,14,12,8,8,8)mm.
When we weld the tank by using submerged arc welding technique (electrode spec. EH12K, DIA 3mm,4mm), we notice distortion in welding joint in bottom 9mm and floating roof 7mm and
shell at thickness of 12,10&8 mm. Please advise me what is the proper electrode and flux to weld this tank without causing distortion in welding joint."
My standard answer is as follows:
"Unfortunately, distortion problems are among the most difficult and intractable of welding practice. Contrary to what you may think they will not go out by a simple change of electrode or flux.
You may wish to see my page
In general the problem is attenuated by careful planning of the weld sequence and by limiting heat input."
What this means in practice is that a complete description of the structure must be secured and that a logical sequence must be established to limit concentrating heat input in a limited area.
This analysis is best done by an experienced welding engineer capable to calculate heat input based on used parameters and to exploit fabrication sequences so as to minimize residual stresses. In certain situations it may be possible to straighten out deformed plates or profiles.
See also PWL#104B.
Brazing with Nickel AWS BNi-2/AMS 4777
[From PWL#105 - Section 3]
A quite common question from shops using AWS BNi-2/AMS 4777 nickel base filler brazing alloy to join elevated temperature resistant alloys in vacuum furnaces, refers to the frequent appearance of a center crack right through the brazing material.
Specification for Filler Metals for Brazing and Braze Welding
American Welding Society / 17-Jun-2011 / 62 pages
Click to order.
The brazing range of this filler metal is given as 1010-1175 0C (1850-2150 0F), relatively low, due to melting point depressants included in its composition.
The added elements performing this function (boron, silicon, phosphorus) also reduce the ductility of the brazing alloy. As the solidification occurs in the direction from the outside toward the central part of the brazing gap, and as these elements are the last ones to freeze, it so happens that their concentration is higher in the center.
The location of concentration of those elements is also a weaker point because of their reduced ductility, prone to cracking in service but also, occasionally, during solidification.
If the capillary clearance existing at the brazing temperature results larger than 0.1mm (0.004"), then there are good chances that the brittle constituents solidify at the joint center, visible in a metallographic section as a dark line, indicating a continuous layer likely to crack.
If however the said clearance is smaller than that, the brittle parts solidify preferably as separated islands in a more ductile matrix, capable of sustaining minor strains without cracking.
Therefore, for successful brazing with this brazing alloy, besides all other precautions, it is essential to control the actual capillary gap at brazing temperature, to keep it consistently smaller than 0.05mm (0.002").
SAW of AL6XN
[From PWL#106 - Section 3]
Submerged arc welding (SAW) of superaustenitic stainless steel AL6XN plates 0.325" (9.5 mm) thick, with electrode ERNiCrMo-3 (also called Alloy 625) resulted in frequent appearance of centerline cracking in the root pass. However the same filler metal as covered electrode (ENiCrMo-3) in manual SMAW produced sound welds.
Damian J. Kotecki deals with this query in his note published on page 20 of the Welding Journal of July 2008.
He observes that the above material, whose composition is given by UNS (Unified Numbering System) N08367 where N is the prefix for nickel base alloys, although iron (Fe) is its main component.
He then agrees that the filler metal selected follows common experience that, to avoid pitting corrosion, Ni overmatching must be assured, as provided by the selected filler.
The observed cracks are probably occurring during solidification from the extended temperature range where liquid film is still present while the weld is contracting.
He discusses the roles of dilution and of Niobium (Nb) content. Regarding resistance to solidification cracking, intermediate levels of niobium are dangerous. ERNiCrMo-3 contains 3.15-4.15 %Nb.
According to the author, dangerous niobium levels are more likely to appear with the extensive dilution caused by SAW than with the more limited dilution of SMAW. A similar problem is likely to occur also with 9%Ni steels.
As techniques to reduce dilution (by reducing current) risk to affect productivity adversely, he suggests to switch to other filler metals devoid of niobium, like ERNiCrMo-10 (Alloy 22 or Ni 6022) or ERNiCrNi-4 (Alloy 276 or Ni6276).
Interested readers are referred to the original article mentioned above.
Microfissures in Stainless Steel
[From PWL#107 - Section 3]
It is not the first time that I recommend to the readers the notes that Damian J. Kotecki publishes on the Welding Journal. I believe that a collection of those questions and answers, would be a most instructive reading.
This time I propose to the readers' attention the note published at page 14 on the Welding Journal of November 2008. A worried reader asks about the causes for the appearance of microfissures in the bend tests of welded specimens of type 310 and 330 austenitic stainless steels.
It should be noted that normally the presence of small (less than 0.125" = 3.18 mm) openings is allowed by the bend test requirements. Nevertheless there could be reason for concern.
The Author explains that it is almost impossible to avoid their presence in completely austenitic structures devoid of any ferrite, despite trying to reduce their number by using low heat input and higher purity filler metal.
A suggested alternative to the bend test, is a normal tensile test to be performed on a longitudinal weld specimen and to be stopped at about 10% strain. The article reports on research published in the Welding Research Council Bulletin 502 (www.forengineers.org), which is reassuring, in that also microstructures containing limited microfissures are known to perform well in service.
Interested readers are urged to seek the original article.
Exacting Purity Requirements drive Purging Innovations
[From PWL#108 - Section 3]
Traditional practice required back side purging for certain root welds. This was achieved in the past by simple means assuring protective argon, flow arranged locally.
More recently however, bioprocessing industries, food and pharmaceutical production and semiconductor manufacturing operations, developed increased purity requirements for equipment involved, especially concerning pipe work.
These requirements were formalized in the latest version of:
American Society of Mechanical Engineers / 31-Mar-2011 / 400 pages
Click to Order.
An article introducing the new requirements was published at page 36 of the July 2012 issue of the Welding Journal. The following sentence is enlightening.
"In the food processing industries, statutory legislation and a plethora of litigation suits have forced plant manufacturers to introduce quality control levels previously considered unnecessary. Contamination introduced during fabrication is now unacceptable."
To satisfy these requirements, special purging devices were developed, essentially delimiting a certain volume in the pipe where welding has to be performed. This is achieved by using specially made expanding plugs, through which back face flowing shielding gas is provided.
Heat resisting materials are employed for those cases where preheat or post weld heat treatment operations are required. With such implements it becomes possible to meet the new specification demands.
Interested readers are urged to seek the original article indicated above.
How to develop a laser welding procedure
[From PWL#109 -Section 3]
Question: I have a small 11.5 gauge T-304 SS Tubing and need to weld a clean solid bead all the way around to a 17-4 SS Insert with a 100W Co2 Laser. What would you recommend for setting e.g.: power, velocity, PWM (Pulse-width modulation) Frequency, cover gas, filler?
Answer: Unfortunately the question above cannot be answered to give a real, practical procedure applicable to the set-up indicated above, or for that matter to any set-up at all.
It should be noted that the joint configuration is not described in sufficient detail and that no requirements are spelled out to convey the important elements for quality evaluation of results.
The request of specific parameter values for process variables, as if a single solution were possible, ignores the fact that the actual selections influence each other in subtle ways: it indicates that the inquirer is not alert to the practical aspects of the development process.
Even if a numerical simulation program was available providing tentative settings for starting trials, nonetheless physical test runs would be required to permit refinement and fine tuning of the real values.
One should be aware of the empiric fact that different combinations of independent process variables or welding parameters may be found to give acceptable results. A thorough development program should explore most of the possible procedures, among which the most suitable would then be selected.
Also one should examine if Laser Beam Welding (LBW) is indeed the most suitable process for the job or if the choice was only dictated by the availability of this equipment.
The right way to find a suitable answer to the question above is to set a development program to be carried out on scrap material, and to be prepared to change the parameters by small amounts, in sequence, until an acceptable weld is obtained.
A thorough treatment of this subject can be found at page 556 of ASM Handbook Volume 6A under the title "Laser Beam Welding".
The article above (14 pages) can be purchased online from
Price $30.00, Member Price $24.00
The first parameter to be selected is the actual power density, depending not only on the equipment output, but also on beam diameter and its spatial distribution. Also position and depth of focus relative to the workpiece surface affect the value obtained.
The equipment may provide pulsed or continuous wave laser. Depending on the welding results, the tentative selection of a definite mode may indicate which one seems to be more promising.
Previous experience with the equipment could help in selecting a tentative power value (maybe around half of the maximum power available) that will give an acceptable static or low speed weld bead without too much spatter and without burning through.
Interaction time describes the time a certain spot of the workpiece is under the laser beam.
The speed selected affects this time, the weld profile and the depth of beam penetration.
The interplay between power and speed determines weld shape, penetration and microstructure, that will be evaluated by performing metallographic examinations of cross sections through the weld.
During this exercise one could play around with focal spot and beam diameter, to find suitable starting parameters. Power could then be raised gradually until acceptable values are found. Then the total heat input can be decreased if desired, by introducing pulse-width modulation.
If filler metal is required, type 308L could be selected and preplaced on the joint. Preliminary welding tests can be performed in air. Production welds can be protected from discoloration by a stream of low pressure flowing argon directed to cover the weld area.
Once the visual results indicate that a suitable procedure is reached, one should perform and check if all requirements are met. The outcome of the tests may require further refinements of the procedure.
Welding Bolts made of ASTM A320 grade L7
[From PWL#110 - Section 3]
Question: What are the problems when welding bolts made of ASTM A320 grade L7?
Answer: This material is intended for low-temperature service down to -150°F (-101°C) and has a minimum Charpy impact value of 15 ft-lb at this temperature. Sizes 2-1/2 in. and under. Bolts are heat treated to Tensile Strength 125 ksi min (860 MPa min), Yield Strength 105 ksi min (725 MPa min), Elongation 16% min, Reduction of Area 50% min.
Welding this bolt material will affect mechanical properties: depending on applications one should make sure that suitability is not compromised.
From the large range of carbon content (0.38-0.48 %C) one can see why three AISI steels, classified as chrome- molybdenum steels, are reported as equivalents, namely 4140, 4142, 4145.
The higher the carbon the higher the susceptibility to hardening and to cracking.
Welding requires the use of low hydrogen electrodes, suitable preheating to reduce risk of cracking, and determination if properties after welding are sufficient for the purpose.
A summarized yet thorough treatment of Hydrogen Induced Cracking was given (4) in PWL#044. Click on PWL#044 to see it.
A practical approach to the problem is presented in the Hydrogen Embrittlement website page.
If mechanical properties after welding are insufficient for the purpose, then full re-heat treatment may be required.
For establishing correct heat treating procedures, one should first perform chemical analysis of the material batch in question, to calculate the carbon equivalent (see http://www.welding-advisers.com/Welding-alloy-steel.html).
Precautions should be taken to avoid decarburization during heat treatment.
Guided Bend Test for welded Aluminum
[From PWL#111 - Section 3]
In an interesting Q&A Note, published on the October 2012 issue of the Welding Journal at page 18, Tony Anderson explains basic requirements for bend test.
That follows an extensive discussion, reported hereafter in section 4, on how to select filler metals for application in welding aluminum alloys.
The Author explains that the usual plunge type fixture used for steel bend test specimens, is not recommended for testing welded aluminum specimens.
That is because, due to the properties of the aluminum heat affected zone, using the plunger may cause the specimen to bend sharply and break.
Instead of that, a suitable wraparound guided bend test fixture forces the test specimen to bend progressively around a pin. All parts of weld and heat affected zone are bent around the same radius of curvature and, therefore, are submitted to the same strain level.
The Author stresses also two other most important points. One is the maximum radius permitted by codes on the corners of the specimen. That is usually up to 3 mm (0.125"), and should be adhered to, as sharp corners invite failure.
The other is the correct bend radius stipulated by AWS D1.2 - Structural Welding Code - Aluminum that varies with materials and condition (as welded or annealed), and the specimen thickness, that should be as required before performing the bend test. Annealing, if required, should be performed as specified in the code.
Readers facing problems in meeting bend test requirements in welded aluminum specimens, are urged to seek the original article quoted above.
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