This effort has seen support from the government and within the steel industry due to its potential for cutting down energy use and environmental emissions, particularly carbon dioxide. The process utilizes direct gaseous reduction of fine iron concentrates to make iron and is being powered by natural gas. Berry’s role in the project so far, and in the future, will be bringing the technology to life through designing and manufacturing the equipment, enabling practical application.
IPC Crowns Winner of Hand Soldering World Championship
attracted participants from around the globe. Fu Chunyan of Beijing Railway Signal Co. Ltd. in China earned first place. She won $1000; a new Metcal soldering station; a Mantis inspection system; and a Sovella ESD footrest. In addition, second place went to Viengkeo (Gail) Sourivongs of Connecticut-based Imperial Electronic Assembly, and the third place prize went to Wang He of China’s Changchun Institute of Optics. The contestants built a functional electronics assembly within a 1-h time limit that was judged in accordance with IPC-A-610E Class 3 criteria, production speed, and overall electrical functionality. Upcoming hand soldering competitions are also planned for Malaysia, Thailand, and India. For more details, visit www.ipc.org/hsc.
Coxreels Celebrates Its 90th Anniversary
Fu Chunyan of Bei jing Railway Signal Co. Ltd. accepts her fi rst place IPC World Championship hand soldering competition award from John Mitchell, IPC president and CEO.
The first IPC World Championship hand soldering competition at IPC APEX EXPO® in San Diego, Calif., on February 21,
Coxreels, Tempe, Ariz., a manufacturer of hose, cord, and cable reels, is celebrating its 90th anniversary this year. The thirdgeneration, family owned and operated business was established in 1923 as Cox Air Gauge. Originally aimed at enhancing the automotive service station market, its offering has grown into a global product used in more than 24 industries. Many patents have marked the company’s milestones. It also designs, builds, and supports all its products in the United States. Coxreels has been in business for 90 years. Displayed is one of the company’s oldest known catalogs from when it was known as Coxwells Inc.
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Industry Notes • Camfil Farr Air Pollution Control (APC) , Jonesboro, Ark., a producer of dust and fume collectors to clean up industrial processes, plus an America n Welding Society and International Thermal Spray Association member, will now operate as Camfil Air Pollution Control (APC) . The new Web address for the company is www.camfilapc.com.
• Masterweld Products, Coplay, Pa., a provider of gas metal arc, gas tungsten arc, and plasma torches as well as consumables, has expanded its repair department to include the Houston, Tex., location. The facility will begin operations on May 1.
• Engine manufacturers, including KOHLER Engines, are warning users of gas-powered lawnmowers and other outdoor power equipment to be vigilant when fueling. Blends with more than 10% ethanol, such as E15 and E85, should not be used. They can cause permanent, irreversible damage not covered under warranty.
• The Aluminum Association, Arlington, Va., and Metal Powder Industries Federation , Princeton, N.J., signed a memorandum
of understanding to share metal powder safety information.
• Goodwill Industries of Northwest North Carolina and Forsyth Technical Community College, Winston-Salem, N.C., have
teamed up to offer welding classes for skills sought by area companies such as Siemens, Caterpillar, and John DeereHitachi. For more details, visit www.goodwillclasses.org .
• Desert NDT, LLC, Odessa, Tex., has acquired T&K Inspection, Inc., Williston, N.D. Co-owners Jerry Thompson and Ken Kain
will continue to oversee loc al office operations.◆
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WELDING JOURNAL
11
INTERNATIONAL UPDATE
New Facility Supports Solid-State Laser Development
the sales organization and customers with technical training seminars and detailed research into material applications. It will also provide materials analysis for other Sandvik business areas, specifically Sandvik Construction and Sandvik Machining Solutions.
Wall Colmonoy Joins Forces with Metaglobal
The TRUMPF Development Center, with a total floor space of 6200 sq m, contains office space and development laboratories for solid-state lasers.
TRUMPF, a manufacturer of lasers for industrial use, has expanded its primary solid-state laser development facility by erecting a new building in the town of Schramberg-Sulgen, Germany. TRUMPF Laser GmbH + Co. KG has augmented this corporate site with a structure offering 6200 sq m of floor space. The twostory structure, measuring 52 × 52 m, houses laser testing laboratories, a climate control chamber, and the building’s utility services on the first floor. The upper floor consists of offices and conference rooms. With an investment of $17.5 million and 17 months of construction work, development sections at Schramberg, which were previously located in a number of different buildings around the site, are now consolidated under a single roof. The company plans to use the freed up floor space to expand its production capacities for solid-state lasers.
Wall Colmonoy’s European headquarters in Pontardawe, Wales, has reached a distributor partnership agreement with Portugalbased supplier, Metaglobal LDA, to sell surfacing and brazing products, and castings to the Iberian region. The agreement will strengthen Wall Colmonoy’s presence, market coverage, and support to customers in the Iberian region. Metaglobal LDA was founded in 2006, and has facilities in Lisbon and Leiria, Portugal. Its sales offices and staff will offer customers support on products, services, and technologies in industries such as glass, oil and gas, automotive, and aerospace. Richard Shaw, commercial director of Wall Colmonoy Ltd., said, “We are excited to partner with a distributor that can expand our product reach and deliver technical support that customers can rely on.” Certain global accounts will continue to be managed directly by Wall Colmonoy Ltd.; however, all new inquiries for Spain and Portugal will be handled by Metaglobal.
Westinghouse to Support Argentinaʼs Steam Generator Replacement
Chinese Research Center to Feature Specialized Laboratories
Westinghouse will provide welding services in support of project Life Extension Embalse Nucle ar Power Plant. (Photo courtesy of Nucleoeléctrica S. A.) The Sandvik Materials Technology management team and Zhenjiang local government officials turned the first soil on the site of the new research center.
Sandvik Materials Technology, a global engineering company, plans to invest in a new high-tech research and development center adjacent to its manufacturing facilities in Zhenjiang, China. A symbolic ground-breaking ceremony was held on March 6. Building work on the 1440-sq-m center is set to begin this summer, with plans to be operational in early 2014. It will accommodate specialized laboratories, a learning center, offices, and an exhibition area. ZZ Zhang, Sandvik China president, said, “The laboratories will be equipped to the highest standards with modern analytical equipment, including scanning electron microscopes and advanced mechanical testing facilities.” The center will support not only production units, but also 12
MAY 2013
Westinghouse Electric Co. recently announced that its subsidiary, PCI Energy Services, LLC (PCI), has signed a contract with Nucleoeléctrica Argentina S.A. (NA-SA) to provide engineering, specialty pipe cutting, and welding services in support of the replacement steam generator program at Argentina’s Embalse Nuclear Power Plant. This work is part of the overall refurbishment program at Embalse to extend the plant’s life by up to an additional 30 years. Rubén Semmoloni, NA-SA project director, Life Extension Embalse Nuclear Power Plant, said, “NA-SA values the engagement and participation of PCI in the Embalse life extension project, which will further enhance plant safety and operation for years to come.” Although the engineering scope of the work for the CANDU-6 pressurized heavy water reactor plant is underway, the major site activities are expected to be executed during 2014.
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STAINLESS Q&A
BY DAMIAN J. KOTECKI
Q: We have fabricated some 304 stainless steel tanks using ER308LSi filler metal. Our customer specified the 304, rather than 304L, out of cost concerns. Ordinary tap water is corroding the tanks, beside the welds. One corroded tank was examined and found to be sensitized in the heat-affected zone (HAZ), where the corrosion is occurring. Can we do anything about the remaining tanks?
A: Sensitization was previously addressed in the November 2007 Stainless Q&A column. Briefly, sensitization occurs in those portions of the heat-affected zone (HAZ) that reach a peak temperature between about 480° and 870°C (900° and 1600°F) when there is enough carbon (more than 0.03%) available to produce precipitation of chromium-rich carbides along grain boundaries. Higher peak temperatures than 870°C either allow chromium to diffuse fast enough to keep up with the carbon in forming carbides, or actually cause the carbides to dissolve. Peak temperatures below about 480°C don’t allow enough carbon diffusion to form significant chromium carbides during welding. The carbides have the general formula M23C6, where M is any metallic element, but chromium is by far the most concentrated metallic element in the carbides. The carbon atom is a very small atom that can diffuse rapidly through the stainless steel matrix to the grain boundaries, so that carbon from anywhere in a grain can reach the grain boundary in this temperature range. But the chromium atom is a large atom that diffuses slowly, so that only chromium from very close to the grain boundary participates in formation of the carbides. Formation of the carbides then tends to produce a chromium-depleted zone beside the grain boundary. This chromium-depleted zone, if exposed to a corrosive medium, is preferentially attacked and dissolved. The corrosion follows the chromium-depleted zones beside the grain boundaries and a continuous network of corrosion along grain boundaries causes grains to separate from the weldment. Figure 1 shows the networks of chromium carbides along the grain boundaries in a 304 HAZ. First, one might ask why your customer would specify 304 instead of 304L. The concern about cost made sense many years ago, but it really doesn’t now. As recently as 1955, the only available methods of producing low-carbon stainless steel involved decarburizing the melt under oxidizing conditions that removed chromium to the slag (Ref. 1). As a result, the melt during 14
MAY 2013
Fig. 1 — Sensitized 304 stainless steel. The chromium carbides responsible for sensitization appear as black specks along the austenite grain boundaries.
Fig. 2 — Graphic display of the effects of time and temperature on chromium carbide pre cipitation and intergranular corrosion (IC) in 304 stainless steel.
decarburizing was designed to contain only about 2% Cr. Then after decarburization, expensive low-carbon ferrochromium was added to the melt to reach the intended chromium content in the stainless steel. But around 1955, Dr.
William A. Krivsky and his colleagues at what was then the Linde Division of Union Carbide Corp. developed the argonoxygen decarburization (AOD) process. In the AOD process, much cheaper high-carbon ferrochromium was intro-
duced into the melt and an argon-oxygen gas mixture was forced into the melt. Combustion of the carbon raised the temperature of the melt and removed the carbon. As the carbon decreased, the oxygen content of the gas stream was decreased. Then ferrosilicon was added to recover some oxidized chromium from the slag. The decarburization could be stopped at whatever carbon level was desired. The argon cost was more than offset by using cheap high-carbon ferrochromium instead of expensive low-carbon ferrochromium for the alloying. It took about 15 years to fully debug the AOD process, but since about 1970, the AOD process has been the preferred method of stainless steel refining. It completely dominates stainless steel refining in the Western World. Today, I am told that the price premium for 304L over 304 is about 1.5 to 2 cents/lb. That price difference is hardly worth considering in view of the likelihood of sensitization when welding 304. There is a second reason that 304 might be specified. Its specified minimum tensile strength is 75 ksi (515 MPa) vs. 70 ksi (485 MPa) for 304L, and the specified minimum yield strength of 304 is 30 ksi (205 MPa) vs. 25 ksi (170 MPa) according to ASTM A240. However, you can purchase stainless steel that is dual certified as both 304 and 304L (i.e., it is below 0.03% carbon but still meets the higher strength requirement of 304). That would take care of the strength concern. The above discussion does not solve your problem, but hopefully it will keep others from making the same mistake. Once 304 (or any other nonlow-carbon stainless steel) has been sensitized, there are only two ways of removing the sensitization; both involve heat treatment, and neither is very palatable. The first is a solution anneal followed immediately by water quench. The annealing temperature of about 1040°C (1900°F), for an hour or so, dissolves all of the chromium carbides and diffuses chromium back into the chromium-depleted zones beside the grain boundaries. The water quench from the annealing temperature is necessary to keep the carbon in solution. The problem with this approach, however, besides cost, is that the stainless steel oxidizes heavily at the annealing temperature and tends to distort both at the annealing temperature and during the quench. A cylindrical shape, such as a pipe, lends itself to this approach because the cylindrical shape is quite stiff, and entering the quench from one end of an open cylinder greatly limits distortion. But a tank does not lend itself readily to this approach. If you were to anneal and quench your tanks, I believe you would need to rigidly support the tanks to maintain their present shape, and you would have to descale after the quench.
The second heat treatment involves considerably lower temperature, but much longer time. The idea behind this second heat treatment is to allow the carbides to form, but to provide enough time at temperature to also allow chromium to diffuse back into the chromium-depleted zones and eliminate them. This can require more than 100 h at a temperature like 760°C (1400°F). Figure 2 illustrates the effect of temperature on chromium carbide precipitation and on intergranular corrosion (labeled “IC attack” in the figure). It can be seen from this figure, reproduced from Folkhard (Ref. 2), that long times at intermediate temperatures can be used to heal the damage from sensitization. The advantages of this approach over annealing and quenching is that distortion will be much less (in part because no quench is required at the end of the treatment) and scaling will not occur although the steel will be oxidized. So cleanup afterward is less. But 100 h at 760°C is not cheap. In conclusion, use of 304 in a weldment that will see corrosive service is not a good idea. The fix is expensive. ♦
References
1. Krivsky, W. A. 1973. Stainless History, Metallurgical and Materials Transactions, Vol. 4, No. 6, pp. 1439–1477. 2. Folkhard, E. 1988. Welding Metal lurgy of Stainless Steels . Springer-Verlag, Vienna.
DAMIAN J. KOTECKI is president, Damian Kotecki Welding Consultants, Inc. He is treasurer of the IIW and a member of the A5D Subcommittee on Stainless Steel Filler Metals, D1K Subcommittee on Stain less Steel Structural Welding; and WRC Subcommittee on Welding Stainless Steels and Nickel-Base Alloys. He is a past chair of the A5 Committee on Filler Metals and Al lied Materials, and served as AWS president (2005–2006). Send questions to damian@ damiankotecki.com, or Damian Kotecki, c/o Welding Journal Dept., 8669 Doral Blvd., Ste. 130, Doral, FL 33166.
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WELDING JOURNAL
15
Friends and Colleagues:
I want to encourage you to submit nomination packages for those individuals whom you feel have a history of accomplishments and contributions to our profession consistent with the standards set by the existing Fellows. In particular, I would make a special request that you look to the most senior members of your Section or District in considering members for nomination. In many cases, the colleagues and peers of these individuals who are the most familiar with thei r contributions, and who would normally nominate the candidate, are no longer with us. I want to be sure that we take the extra effort required to make sure that those truly worthy are not overlooked because no obvious individual was available to start the nomination process. For specifics on the nomination requirements, please contact Wendy Sue Reeve at AWS headquarters in Miami, or simply follow the instructions on the Fellow nomination form in this issue of the Welding Journal. Please remember, we all benefit in the honoring of those who have made major contributions to our chosen profession and livelihood. The deadline for submission is July 1, 2013. The Committee looks forward to receiving numerous Fellow nominations for 2014 consideration.
Sincerely, Thomas M. Mustaleski Chair, AWS Fellows Selection Committee
Fellow Description DEFINITION AND HISTORY The American Welding Society, in 1990, established the honor of Fellow of the Society to recognize members for distinguished contributions to the field of welding science and technology, and for promoting and sustaining the professional stature of the field. Election as a Fellow of the Society is based on the outstanding accomplishments and technical impact of the individual. Such accomplishments will have advanced the science, technology and application of welding, as evidenced by: ∗ Sustained service and performance in the advancement of welding science and technology ∗ Publication of papers, articles and books which enhance knowledge of welding ∗ Innovative development of welding technology ∗ Society and chapter contributions ∗ Professional recognition RULES 1. 2. 3. 4. 5. 6. 7.
Candidates shall have 10 years of membership in AWS Candidates shall be nominated by any five members of the Society Nominations shall be submitted on the official form available from AWS Headquarters Nominations must be submitted to AWS Headquarters no later than July 1 of the year prior to that in which the award is to be presented Nominations will remain valid for three years All information on nominees will be held in strict confidence No more than two posthumous Fellows may be elected each year
NUMBER OF FELLOWS Maximum of 10 Fellows selected each year.
AWS Fellow Application Guidelines Nomination packages for AWS Fellow should clearly demonstrate the candidates outstanding contributions to the advancement of welding science and technology. In order for the Fellows Selection Committee to fairly assess the candidates qualifications, the nomination package must list and clearly describe the candidates specific technical accomplishments, how they contributed to the advancement of welding technology, and that these contributions were sustained. Essential in demonstrating the candidates impact are the following (in approximate order of importance). 1. 2. 3. 4. 5. 6. 7. 8. 9. 10. 11.
Description of significant technical advancements. This should be a brief summary of the candidates most significant contributions to the advancement of welding science and technology. Publications of books, papers, articles or other significant scholarly works that demonstrate the contributions cited in (1). Where possible, papers and articles should be designated as to whether they were published in peer-reviewed journals. Inventions and patents. Professional recognition including awards and honors from AWS and other professional societies. Meaningful participation in technical committees. Indicate the number of years served on these committees and any leadership roles (chair, vice-chair, subcommittee responsibilities, etc.). Contributions to handbooks and standards. Presentations made at technical conferences and section meetings. Consultancy — particularly as it impacts technology advancement. Leadership at the technical society or corporate level, particularly as it impacts advancement of welding technology. Participation on organizing committees for technical programming. Advocacy — support of the society and its technical advancement through institutional, political or other means.
Note: Application packages that do not support the candidate using the metrics listed above will have a very low probability of success. Supporting Letters Letters of support from individuals knowledgeable of the candidate and his/her contributions are encouraged. These letters should address the metrics listed above and provide personal insight into the contributions and stature of the candidate. Letters of support that simply endorse the candidate will have little impact on the selection process. Return completed Fellow nomination package to: Wendy S. Reeve American Welding Society Senior Manager Award Programs and Administrative Support 8669 Doral Blvd., Suite 130 Doral, FL 33166 Telephone: 800-443-9353, extension 293
SUBMISSION DEADLINE: July 1, 2013
(please type or print in black ink) CLASS OF 2014 FELLOW NOMINATION FORM
DATE_________________NAME OF CANDIDATE________________________________________________________________________ AWS MEMBER NO.___________________________YEARS OF AWS MEMBERSHIP____________________________________________ HOME ADDRESS____________________________________________________________________________________________________ CITY_______________________________________________STATE________ZIP CODE__________PHONE________________________ PRESENT COMPANY/INSTITUTION AFFILIATION_______________________________________________________________________ TITLE/POSITION____________________________________________________________________________________________________ BUSINESS ADDRESS________________________________________________________________________________________________ CITY______________________________________________STATE________ZIP CODE__________PHONE_________________________ ACADEMIC BACKGROUND, AS APPLICABLE: INSTITUTION______________________________________________________________________________________________________ MAJOR & MINOR__________________________________________________________________________________________________ DEGREES OR CERTIFICATES/YEAR____________________________________________________________________________________ LICENSED PROFESSIONAL ENGINEER: YES_________NO__________ STATE______________________________________________ SIGNIFICANT WORK EXPERIENCE: COMPANY/CITY/STATE_____________________________________________________________________________________________ POSITION____________________________________________________________________________YEARS_______________________ COMPANY/CITY/STATE_____________________________________________________________________________________________ POSITION____________________________________________________________________________YEARS_______________________ SUMMARIZE MAJOR CONTRIBUTIONS IN THESE POSITIONS: __________________________________________________________________________________________________________________ __________________________________________________________________________________________________________________ __________________________________________________________________________________________________________________ IT IS MANDATORY THAT A CITATION (50 TO 100 WORDS, USE SEPARATE SHEET) INDICATING WHY THE NOMINEE SHOULD BE SELECTED AS AN AWS FELLOW ACCOMPANY NOMINATION PACKET. IF NOMINEE IS SELECTED, THIS STATEMENT MAY BE INCORPORATED WITHIN THE CITATION CERTIFICATE. SEE GUIDELINES ON REVERSE SIDE
SUBMITTED BY: PROPOSER_______________________________________________AWS Member No.___________________ Print Name___________________________________ The Proposer will serve as the contact if the Selection Committee requires further information. Signatures on this nominating form, or supporting letters from each nominator, are required from four AWS members in addition to the Proposer. Signatures may be acquired by photocopying the original and transmitting to each nominating member. Once the signatures are secured, the total package should be submitted. NOMINATING MEMBER:___________________________________NOMINATING MEMBER:___________________________________ Print Name___________________________________ Print Name___________________________________ AWS Member No.______________ AWS Member No.______________ NOMINATING MEMBER:___________________________________NOMINATING MEMBER:___________________________________ Print Name___________________________________ Print Name___________________________________ AWS Member No.______________ AWS Member No.______________
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RWMA Q&A Q: I am trying to make resistance welding seams using a single-phase constant current welding control and am having a hard time holding the tolerance required for this military project. We are using a 3 8150-kVA seam welding machine with ⁄ in.-wide welding wheels on 0.040-in. CRS. The welding transformer tap switch is set to the #1 position. I checked the learn
Fig. 1 — 99% weld heat.
BY ROGER HIRSCH
table in the control and see that we are in the 25–30% range so I know I am not over worki ng the weldi ng machi ne. Do you have any suggestions?
A :
The problem here is a misunderstanding of how a resistance welding machine works. Because you are using the control in this very low heat percentage
Fig. 2 — 50% weld heat.
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range, the output of the welding transformer is a series of very small heat pulses and a lot of spaces in between. This makes control of the process very sensitive. No matter how good your control is, it will be very difficult to achieve the desired results with this welding machine. Welding machine size is also often misunderstood. The idea that you need a
Fig. 3 — 30% weld heat.
welding machine large enough to join the thickest metal can also mean it will be way too large to do the smaller thickness combinations. Some of this problem can be overcome by using the welding machine with the transformer tap switch set at the high settings for the thicker metal welding, and then going to the low tap switch setting for thin metals. Sadly, many of the newer U.S.-made welding machines and most of the imported welding machines do not have a transformer tap switch and, therefore, lose this ability. To understand why using a welding machine at a very low heat percent setting is a problem, you have to understand how a welding control provides heat in response to the program settings. The process is called phase shifting . This is accomplished by having the welding control fire the SCR contactor (solid-state switch that conducts voltage to the welding transformer) at different time delays in each half cycle of the line power. If you want to use all of the line voltage, the control will turn on the SCR contactor just after the current in the welding transformer goes to zero. Since a welding transformer is an inductive device, this will happen a little after the line voltage goes to zero. Figure 1 shows a welding machine operating at a 99% heat setting. The lower red trace shows the current going into the welding transformer primary. Note that the current is conducted over the entire sine wave of line power less a small notch at each zero crossing. This would be about the same as if you shorted the SCR contactor and put the entire available line voltage into the welding machine transformer. The upper blue scan shows the RMS current created by this AC firing. This RMS current is proportional to the output of the welding transformer. Figure 2 shows a current scan for a heat setting of 50%. Note that about half of the sine wave is being used, and the other half
Fig. 4 — 460-V welding machine operating on a 230-V line at 60% weld heat.
has no heat. The upper scan shows the RMS current that results from this firing. Figure 3 shows what happens with a heat setting of 30%, which is similar to your setup. Note that the amount of time that voltage is being conducted to the welding transformer is very small compared to the time of no voltage flow in
each half cycle. The upper scan shows the RMS current from this setting. A very small change in this heat setting makes a large change in the RMS current, and a control working in this range will be unstable and cannot be accurate. A good rule of thumb is to use a resistance welding machine with a transformer
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WELDING JOURNAL
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tap switch setting that will produce the desired RMS welding current when the control’s heat setting is between 60 and 90%. The recommended upper limit of 90% is to allow the welding control to have the ability to increase heat to compensate for incoming line voltage variations. It also allows for changes in the welding machine secondary impedance as ferrous metal of the part is being pushed into the welding machine throat. There are several solutions to this type of control problem: The first is to set the welding machine transformer tap switch to a lower number and then use a higher weld heat setting. Unfortunately, in your case, you are already at the lowest setting. Next, use the correct size welding machine for the job. If you have a smaller kVA seam welding machine, move the job there. As an example, a 75-kVA seam welding machine will probably find the proper heat setting in the 60% range to give control back to the system. If this is not possible, and if you are working on a 460-V power line, co nnect the welding machine to a 230-V line with the results as shown in Fig. 4. Be sure to change the voltage select jumpers in the control to 230 V. Now you will be in the 60% range since the line voltage amplitude will be lower and you will be using more of each ½ cycle of this line voltage to create the desired RMS current for welding. This will not increase the load on the power lines since the turns ratio of the welding transformer will remain the same. Note that you can operate a 460-V transformer on a 230-V line, but you cannot operate a 230-V transformer on a 460V line without damaging the transformer. Compare this to Fig. 3, which shows the same transformer operating on 460 V at a 30% weld heat setting. You can easily see how much more stable the process is when more of each half-cycle of current is being used.
form a nugget. When you try to make the second spot weld, some of the voltage being conducted from the upper to the lower electrode finds an easy path through the fused metal of the first weld nugget. This is called shunting . Because of shunting, the second weld will be considerably weaker than the first weld. When you do the third and fourth welds, the same problem continues, but at an even greater level of strength loss. Making four welds this close together will, if you are lucky, produce a total 1 2 times that of a single strength of about 1 ⁄ weld. And since the diameter of each weld nugget will be small enough to fit into this tab, the total strength of all four welds will be very low as you have observed. The solution is to use welding projections on the tabs. This will allow the welding current to make good separate welds at each projection in one pass of the welding machine, and the strength of the overall joint will be much closer to the strength of a single projection multiplied by the number being used. In this case, since you are welding a flat tab to a curved surface, it would be best to use two oval-shaped projections placed at right angles to the radius of the part. This will allow the maximum weld area
and compensate for any slight misalignment of the parts. Another big advantage to using pro jections on this part is that you can use flat electrodes. These electrodes will have much longer life than a small spot welding electrode, and weld strength will not be dependent on how well the electrodes are dressed. ◆
ROGER HIRSCH is past chair of the RWMA, a standing committee of the American Welding Society. He is also president of Unitrol Electronics, Inc., Northbrook, Ill., a manufacturer of re sistance welding controls and process water chill ers. Send your comments and questions to Roger Hirsch at
[email protected] , or by mail to Roger Hirsch, c/o Welding Journal, 8669 NW Doral Blvd., Suite 130, Doral, FL 33166.
Change of Address? Moving? Make sure delivery of your Welding Journal is not interrupted. Contact the Membership Department with your new address information — (800) 443-9353, ext. 204;
[email protected] .
Q: I have a project to join a small hanger strap to the t op bell of a fire extinguisher. The hanger strap has two flat welding tabs. Each tab has a flat welding area of 1 3 about ⁄ 2 × ⁄ 4 in. The drawing specifies four small spot welds on each tab. I t ried doing this and cannot get very strong welds. What am I doing wrong?
A: The problem here is in the hands of the designer. Many people who design sheetmetal parts do not have a full understanding of how spot welding works and, as a result, specify parts that cannot be successfully welded. This part seems to be such a case. When you make the first of the four small spot welds, you fuse the metal to 22
MAY 2013
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Inconel TM
PRODUCT & PRINT SPOTLIGHT
Wire Feeder Designed for Heavy-Duty GMAW and FCAW The SuitCase X-Treme 12VS voltage-sensing wire feeder has been designed for welding in heavy-duty field applications such as construction, structural steel erection, heavy equipment repair, and mobile fabrication. A wire delivery system eases wire loading and provides minimal resistance in feeding. The drive motor assembly and integrated tachometer ensures the wire feed speed accuracy throughout the day so it performs to precise parameters, whether welding with small-diameter solid wi res (0.023 in.) or 5 64 in.). The wire delivery large-diameter cored wires ( ⁄ system makes it easier to load the 12-in. wire spools and reduces drag on the wire by eliminating the inlet guide and allowing the wire to roll over the large radius of the drive rolls. It also features a visual scale on the wire pressure knob and allows welders to dial in tension settings. Other features include a redesigned placement of the shielding gas inlet (for GMAW and dual-shielded FCAW applications) to better protect the fitting from damage, and a wire speed dual schedule feature that reduces wire feed speed to 87.5% of standard speed; this feature requires a dual schedule gun or switch (sold separately). Additional highlights include SunVision™ digital meters; a portable polypropylene case with built-in slide rails and ability to open the door to change wire in a vertical position; and potted/trayed main printed circuit boards. It is compatible with CC or CV DC power sources or engine-driven welding machines/generators. Miller Electric Mfg. Co. www.millerwelds.com (800) 426-4553
System Transforms GMA Gun into Stud Welding Tool
hammer, 100 2-mm pins, and a replacement slide hammer knurled locking cam. The Eastwood Co. www.eastwood.com (800) 343-9353
GMAW Gun Handle Reduces Strain on Shoulders, Arms
The company’s GMA stud weld system transforms any Tweco®-style GMA g un into a stud welding tool, which enables users to remove dings in doors, hoods, deck lids, or quarter panels. After the pins are welded into place using the product, pull the metal flush with the slide hammer, cut and grind the pins, and a properly shaped surface ready to finish is achieved. It includes a GMA stud nozzle attachment, precision-balanced slide 24
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gonomics have been developed and tested with professional welders and the Finnish Institute of Occupational Health. When a welder uses the pistol grip with the handle, its cable passes above the wrist and its center of gravity is closer to the elbow. When welding with the product in a horizontal or vertical position, the upper arm is lower. The pistol grip does not require any squeezing of the handle; it allows a relaxed hold. The cable passing over the lower arm balances the handle. It is compatible with most Abicor Binzel®, Tbi®, and Kemppi® torches. Ergowelder Oy Ltd. www.ergowelder.com +358 (0)44 2779953
Catalog Showcases Line of Women’s Welding Gear The company’s GMAW gun handle allows two working positions, traditional and pistol grip. It also reduces strain on shoulders and arms. The handle’s er-
The 2013 Welding Gear catalog introduces the new Jessi Combs women’s line, which includes the VIKING™ 1840 Series
Amp Angel™ autodarkening welding helmet, Women’s Shadow™ welding jacket, and several glove options sized for female hands. New offerings include custom embroidery, a new facial protection category, welding brushes, welding curtains, and welding blankets. The catalog’s latest edition may be requested from the contact information below.
Welding Machine Useful for Shipyards
Two-Gas Adjustable Mixer Configured for Hydrogen
The portable VR 5000 case has been developed for dusty, damp, and salty en vironments. Used in combination with GMA power sources from the TransSteel series for separate wire feed units, it gives welders a system for shipbuilding, oil-rig construction, railway-vehicle manufacturing, and site erection. With external dimensions of 507 200 320 mm, it fits through any manhole of up to 350 mm in diameter. The unit weighs less than 22 lb and has interconnecting hose packs of up to 229 ft long (for gas-cooled welding systems). It is designed for use with an 11-lb wire spool 200 mm in diameter. The feeder unit is offered in water- or gascooled, or synergic versions.
The company’s SuperFlash Mini-PGM t two-gas adjustable mixer can now be configured for use with hydrogen. Users can customize hydrogen/argon or hydrogen/ nitrogen mixes to best support their process, creating their own mixed gas for GMAW, GTAW, or plasma gouging/cutting/welding. Weighing 7.5 lb and using less than 1 ft 3 of space, the Mini-PGM provides enough gas for eight welding machines at approximately 40–50 ft 3 /h per machine. Each mixer is fully adjustable and comes standard with a mounting bracket.
The Lincoln Electric Co.
Fronius International GmbH
SuperFlash
www.lincolnelectric.com (888) 355-3213
www.fronius.com (877) 376-6487
www.oxyfuelsafety.com (888) 327-7306
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WELDING JOURNAL
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Materials Testing Company Updates Web Site
The company has launched an updated Web site with new features that provide easier access to frequently used materials, as well as all of the information and tools found on the previous site. Calibration customers will now find a tabbed section dedicated to the capabilities and services provided by LTI Metrology, the calibration and dimensional inspection division of the company. The Quick Service Form, accessible under the contact tab or with the contact us button, provides a con venient way to request information on materials testing, calibration, quality assurance, billing, pickup, and delivery service. The new site provides easy navigation and access through the homepage to information targeted to specific customers.
x e d n i d a / g r o . s w a . w w w o t o g o f n i r o F
Laboratory Testing, Inc. www.labtesting.com (800) 219-9095
Fluids Eliminate Hazards from Mineral Deposits The company’s new welding chemical products include a heavy-duty antispatter and two glycol-based cooling fluids. The antispatter, which is nonflammable and solvent based, quickly evaporates to pro vide an effective surface coating and eliminates the need to grind or brush the surface after welding. The cooling fluids, designed for use in plasma, GTAW, GMAW, and resistance welding systems, eliminate the hazards of mineral deposits. The coolants also lubricate the pump, liner, and gasket and seal. They are thermally stable and have dielectric properties that make them suitable for plasma and arc welding systems.
x e d n i d a / g r o . s w a . w w w o t o g o f n i r o F
Thermacut, Inc. www.thermacut.com (800) 932-8312
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Earplugs Available in Three Damping Levels The DI Red 25 dB is a CE-certified industrial hearing protection earplug that incorporates DEC sound technology. Available in three damping levels, this universal plug pro vide s acou stic protection in a discrete miniature package and is designed for n oisy en vironments, either at work or leisure. The product allows air to enter the ear, reducing the occlusion effect normally experienced when using earplugs. All of the company’s products are designed to be interchangeable and upgradable to custom fit molds. Crescendo www.crescendo-hearingprotection.com (305) 463-9304
Abrasives Catalog Includes Video Demonstrations The company’s 2013–2014 Time Saving Solutions catalog features nonwoven cotton-fiber abrasive grinding wheels, cut-
Manufacturing Manufacturin uring
Flux Cored Welding elding Wire W off wheels, mounted points, and other types of grinding, blending, deburring, and finishing products. The catalog includes usage tips; easy-to-read charts with sizes, shapes, and specifications; and QR codes to video demonstrations. New products include Type 27 Max Flex cotton-fiber wheels and rubber-mounted points. The catalog can be requested from the information below or can be downloaded from the Web site.
COBALT NICKEL HARDFACE E STAINLESS S INLESS
Rex-Cut Abrasives www.rexcut.com (800) 225-8182
BELLINGHAM TECHNICAL COLLEGE BELLINGHAM TECHNICAL COLLEGE
WELDING SCULPTURE THEME: SMALL SMALL WORKS WORKS
ALLOY ALLO Y STEEL EEL
w w w . b t c . c t c . e d u
TOOL STEEL STEE EEL MAINTENANCE MAINTENAN CE FORGE ALLO ALLOYS OYS CUSTOM ALLOYS OYS
MAY 17 & 18, BTC CAMPUS 8:0O AMTO 5:00 PM
x e d n i d a / g r o . s w a . w w w o t o g o f n i r o F
WELDED SCULPTURE COMPETITION (By Invitation) SKILLS CHALLENGE (Open to All -- High School Students Encouraged) AUCTION VENDOR & TRADE EXPO FOOD | FUN | FREE
COR-MET, COR -MET,, INC. x e d n i d a / g r o . s w a . w w w o t o g o f n i r o F
12500 Grand and River Rd. Brighton, MI 48116 481 116 PH: 800 800-848-2719 -848-27 719 FAX:: 810 810-227-9266 -227-9266 www.cor-met.com www.cor -met.com t.com
[email protected] sales@cor -met.co t.com WELDING JOURNAL
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Grinders Feature High Power-to-Weight Ratio
The company’s line of CP3850 industrial angle grinders and sanders offers a high power-to-weight ratio with a 2.8-hp motor. The CP3850, available in both 7and 9-in.-capacity models, weighs less than 6 lb, allowing easier handling. Many ergonomic features have been incorporated to improve operator comfort, safety, and productivity, including a vibrationdamping multiposition side handle, auto balancer, and integrated silencer. Safety features include a double-action safety lever and 270-deg swivel guard to protect the operator from debris. The series is engineered for aggressive use in contouring, deburring, cutting, and sanding in the metalworking, transformation, manufacturing, and energy industries.
x e d n i d a / g r o . s w a . w w w o t o g o f n i r o F
Chicago Pneumatic www.cp.com (800) 624-4735
Magnetic Boards Track Lean Initiative Projects Teams focusing on productivity, process, quality, 5S, and other lean manufacturing improvements use the DoDone StepTracker whiteboard a s a point for project management coordination. Boards come with 6, 12, 25, 42, or 65 project step-stage columns for up to 128 project line items. Titles and project column headings can be custom printed directly
x e d n i d a / g r o . s w a . w w w o t
on the board. It arrives ready to use with magnets and is built to stay like new for a lifetime of daily use with heat-fused printing on porcelain-like steel and whiteboard construction.
o g o f n i r o F
Magnatag® Visible Systems www.magnatag.com/steptracker (800) 624-4154
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MAY 2013
For Info go to to www.aws.org/ad-index www.aws.org/ad-index
BY TOM MYERS TOM MYERS is a senior applications engineer, The Lincoln Electric Co., Cleveland, Ohio, www.lincolnelectric.com.
Not Your Father’s Gas Shielded Flux Cored Electrodes
Fig. 1 — In years past, a limited number of FCAW-G electrode options were offered to fabricators.
FCAW-G consumables have steadily evolved into application-specific productivity workhorses
E
lectrodes for the gas shielded, flux cored arc welding (FCAW-G) process were first developed in the late 1950s. Over the next 40 years or so, manufacturers refined and improved these products, offering a fairly limited line of carbon steel and low-alloy steel electrodes for either all-position welding or flat and horizontal only (i.e., in-position) welding — Fig. 1. During this time, there also were rel-
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atively few variations in the formulations of electrodes within a specific American Welding Society (AWS) classification. In many cases, one particular electrode was intended to be used for a variety of applications. Oftentimes it was intended to be used with either 100% carbon dioxide (CO 2) shielding gas or a mixed argon (Ar)/CO 2 shielding gas. Back then, the general manufacturing philosophy was to develop an electrode
that was in essence a “one-size-fits-all” product. This philosophy was partially successful in meeting the demands of the welding market during this time. However, today’s structural engineers and industrial designers are increasingly specifying higher-strength, lower-weight steels for cost savings and productivity considerations, making these base materials a popular choice in many industries. These new specifications demand the
Fig. 2 — The fast-freezing slag system of all-position FCAW electrodes allows for better outof-position welding capabilities.
need for low-alloy FCAW-G electrodes that produce welds with increased tensile and yield strengths (compared to carbon steel electrodes) for welding these higher-strength steels. Other applications require electrodes that produce welds with improved im pact properties. Generally, electrodes needed to produce welds with low-temperature toughness of at least 20 ft·lbf (27 J) at a test temperature of 0°F (–18°C) or –20°F (–29°C). Some applications now require these same absorbed energy values at temperatures of –40°F (–40°C) or even lower. Similarly, operator demand for just the “right” application-specific, flux cored electrode has steadily increased over the past five to ten years to keep up with a growing desire for increased weld productivity, performance, and quality, not to mention aesthetics. Because of this increased specifica-
tion of new materials, combined with demands for more customized, efficient electrodes, manufacturers have been returning to their R&D drawing boards to develop new gas shielded, flux cored consumables.
Fundamentals and Advantages of FCAW Electrodes Flux cored electrodes were originally developed as a higher productivity extension of shielded metal arc welding (SMAW) electrodes. They are, in fact, like a SMAW electrode turned inside out. They consist of a steel tube (i.e., outer steel sheath) with flux inside the tube or at the electrode’s core, hence the name, “flux core.” Because of this design, the electrode can be wound onto a coil or
spool, and with the use of a wire feeder and welding gun, be fed continuously into the weld joint. Flux cored electrodes fall into two fundamentally different categories: selfshielded, flux cored electrodes (FCAWS) and gas shielded, flux cored electrodes (FCAW-G). Gas shielded, flux cored electrodes incorporate a double shielding system by using an external shielding gas as well as a slag system. The shielding gas is required to protect the arc and molten metal from the atmosphere. It also results in exceptionally smooth arc characteristics, compared to self-shielded electrodes. They use either a rutile slag system or a basic slag system. The rutile system is the most common and is characterized by a smooth arc with complete slag coverage of the weld. The basic slag system, while producing a globular metal WELDING JOURNAL
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Fig. 3 — While the use of a one-sizefits-all electrode for a wide range of applications can deliver adequate arc performance, a single electrode cannot perform well in every application.
transfer and thinner slag coverage, can be more resistant to weld cold cracking. Most FCAW-G electrodes are ideal for all-position welding and all deliver great mechanical properties with high deposition rates. They are used effectively in general shop fabrication, structural steel (including seismic applications), shipbuilding, offshore, pipeline, and other applications. Flux cored arc welding electrodes can be used similarly to SMAW electrodes with a few notable benefits in the process itself. First, SMAW electrodes must be fed manually into a weld joint, making only short welds and resulting in a lot of stop and restart areas in the weld. Restart areas generally have a higher chance of containing a weld defect than any other part of the weld. With the FCAW process, the weld can be made for as long as welders can comfortably reach before having to stop the arc and reposition themselves. This results in fewer restart areas in the weld and, ultimately, fewer chances for weld defects. The FCAW process also has a higher operating factor than the SMAW process (where operating factor (%) equals arc time divided by total fabrication time). It’s also easier to use. It operates at higher current levels, which yields higher deposition rates and higher productivity. Finally, FCAW electrodes have higher electrode efficiency than SMAW electrodes. This means that more of the purchased pounds (kg) of electrode end up as deposited weld metal and less is lost through stubs. Flux cored electrodes, with their slag systems, also have inherent advantages over slagless processes, such as gas metal arc welding (GMAW). The fast-freezing slag system of all-position-classified FCAW electrodes allows for better outof-position welding capability, including vertical and overhead, as the slag helps hold the molten metal against gravity — Fig. 2. Flux cored electrodes produce higher deposition rates when welding out of position than do GMAW consumables. In addition, many in-position-classified FCAW electrodes have good penetration
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characteristics, making them ideal for thicker sections of steel plate. Flux cored arc welding electrodes also handle surface contaminants on steel plate better than solid GMAW electrodes. Not only are deoxidizers present in the outer carbon steel sheath of FCAW electrodes, but deoxidizers, denitrifiers, and scavenger elements are also added to the core elements, while solid GMAW electrodes can only rely on the deoxidizers that are present in the raw green rod material, which is drawn down to make them.
Flux cored arc welding electrodes are considered to be “fabricated” electrodes and, thus, provide a good platform for manufacturing new low-alloy electrodes. The outer sheath on FCAW electrodes — even low-alloy types — are fabricated from types of carbon steels that are either strip-based or green rod-based, both of which are commonly available from steel mills. As such, the core ingredients for various FCAW electrodes then can be altered to produce low-alloy weld deposits with differing mechanical properties. In contrast, low-alloy solid GMAW
Fig. 4 — Some industries, such as pipeline fabrication, often require FCAW-G electrodes that produce welds with a minimum low-temperature toughness of 20 ft-lbf (27 J) at –40°F (–40°C).
wire cannot be fabricated. The final chemistry of the electrode can only be achieved by purchasing it as the raw green rod steel. Low-alloy green rod can be more expensive and difficult to source than carbon-steel green rod.
Concerns with One-SizeFits-All Electrodes The traditional multipurpose approach toward FC AW-G electrodes has proven to be increasingly ineffective over
the years. While the use of a one-sizefits-all electrode for a wide range of applications can deliver adequate arc performance, the reach for a single electrode to perform well in every application is just too broad — Fig. 3. As a result, the arc is never optimized. Why? One electrode used with both 100% CO 2 and mixed gas (i.e., 75% Ar/25% CO2) has to have a fine chemical balance in order to meet the minimum and maximum mechanical property requirements of its AWS classifications with either type of shielding gas. Carbon
dioxide is an active gas, meaning that it actively reacts with some of the electrode’s alloys. Less alloy recovery from the electrode occurs in the weld pool, resulting in a slight decrease in mechanical properties, such as ultimate tensile strength and yield strength. Argon, an inert gas, is nonreactive in the arc. Therefore, the more argon in a mixed gas, the more alloy recovery that occurs in the weld pool. This results in a slight increase in both tensile and yield strengths. Hydrogen levels also play a role in why FCAW-G electrodes are becoming more specific, moving away from the one-sizefits-all design structure of the past. Lower levels of diffusible hydrogen in weld deposits means that such welds will have higher resistance to hydrogen-induced cracking. Welding consumables can be classified with an optional diffusible hydrogen designator. These designators include the letter “H” and a number, which indicate maximum milliliters of diffusible hydrogen per 100 g of weld metal. Most FCAW-G electrodes today meet a diffusible hydrogen rating of H8, with some meeting a very low rating of H4. Some industries, such as shipbuilding/barge building, have increasingly pushed the deposition rate capabilities of all-position FCAW-G electrodes. Generally, when welding in position or with gravity, welders can utilize faster wire feed speed procedures to produce higher deposition rates than they can when welding out of po sition or against gravity. However, because of remote welding locations and limited access to their welding equipment, welders often cannot easily turn up their procedures when they transition from out-of-position to in-position welding. Therefore, they need one set of welding procedures for FCAW-G electrodes that produce maximum deposition rates for out-of-position welding and still produce high deposition rates for in-position welding. Many of the original one-size-fits-all FCAW-G electrodes could only be pushed so far before the slag system would not support the additional molten weld metal. Therefore, new
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FCAW-G electrodes were needed with a different type of slag system. These highdeposition or “HD” electrodes have very fast freezing slag systems that better support higher wire feed speed welding procedures. Additionally, many industries have had increased requirements for weld metal with improved impact properties. The original one-size-fits-all FCAW-G electrodes were designed to produce welds with a m inimum low-temperature toughness of 20 ft-lbf (27 J) at 0°F or 20 ft-lbf (27 J) at –20°F. Some industries, such as offshore and pipeline fabrication (Fig. 4), often require FCAW-G electrodes that produce welds with a minimum low temperature toughness of 20 ft-lbf (27 J) at –40°F. These more stringent requirements have necessitated the need for new FCAW-G electrodes with improved impact properties. In other cases, FCAW-G electrodes have been increasingly used on weldments that must be stress relieved after welding. In general, after postweld heat treatment (PWHT), the tensile and yield strength of the weld drops to a certain degree. When the weld from a one-sizefits-all FCAW-G electrode is stress relieved, you can run the risk of the tensile and yield strengths dropping below the minimum specified levels. Therefore, PWHT applications have led to the need for more specialized FCAW-G electrodes that have an altered chemical formulation. These electrodes are designed to have a minimal drop in tensile and yield strength after stress relief. With such changeable factors as shielding gas, diffusible hydrogen levels, deposition rate needs, and mechanical property requirements, as well as different grades of steel, coming into play in the FCAW-G arena, the scope and versatility of a one-size-fits-all electrode has become increasingly narrower. This challenges manufacturers to meet stringent mechanical properties on a consistent basis with traditional multipurpose FCAW-G electrode design.
New Approach to Design To meet variou s industry-specifi c requirements and operator demands for mechanical properties, performance, and aesthetics, manufacturers now are designing and producing applicationtargeted, next-generation FCAW-G consumables with specific industries in mind. In addition, many of the FCAW-G electrodes today have been designed for use with only one type of shielding gas in 34
MAY 2013
Fig. 5 — Manufacturers must balance three components when they design an FCAW-G electrode — operability, mechanical properties, and diffusible hydrogen levels. order to produce optimum operator appeal and the targeted mechanical properties. They either will be for use with 100% CO2 or a mixed blend, consisting of 75– 85% Ar/balance CO2 (with 75% Ar/25% CO 2 the most popular blend). The required shielding gas is now also incorporated into the electrode’s AWS classification number. For example, the “C” in an E71T-1C classified electrode specifies that it is for use with C O 2 shielding gas, while the “M” in an E71T-1M classified electrode specifies that it is for use with mixed shielding gas. Electrodes that are still designed for use with either type of shielding gas are dual classified, such as “E71T-1C/E71T-1M.” Furthermore, operator appeal of lowalloy FCAW-G electrodes also has improved. A welder can weld with a carbon steel FCAW-G electrode or a low-alloy FCAW-G electrode and not really see a difference in arc performance. This results from the fact that manufacturers have succeeded in coming up with a standard slag system for families of electrodes. Individual electrodes can be modified for different applications by tweaking the alloy formulation in the electrode’s core so that welders and fabricators will see similar operating characteristics, no matter the application.
that produces mechanical properties that are so robust that they tip the scale on operability. Think of the three key components of electrode design as a triangle — Fig. 5. On one side, you have operability. On the second side, you have mechanical properties. On the third side, you have diffusible hydrogen levels. Targeted product development has allowed manufacturers to design “families” of FCAW-G electrodes, each aimed for different applications in specific industry segments. Each family balances the three sides of the design triangle to avoid compromising any one of those components and, thus, delivers a robustly performing electrode.
Application-Specific FCAW Electrodes Today, manufacturers of FCAW electrodes offer broad product lines, with many electrodes designed for specific applications and industries. Examples of more specialized FCAW-G electrodes include ones designed for the following uses: • with one specific type of shielding gas such as UltraCore® 71C and UltraCore® 71A85 from The Lincoln Electric Co. • for higher-strength steels (i.e., 80-, 90-, and 100-ksi minimum tensile strength). • “HD” type for high-deposition, out-ofposition capability (i.e., UltraCore® HD-C and HD-M). • for exceptionally high deposition rates in the flat and horizontal positions. • for improved low-temperature toughness properties. • “SR” type for stress-relieved applications. • for pipe welding applications such as Pipeliner® 81M, 101M, and 111M.
Electrode Design
• for chromium-molybdenum (Cr-Mo) steels.
Producing a successful FCAW-G electrode comes down to balance in the design and manufacture of the electrodes. Manufacturers have worked to develop FCAW-G consumables that consistently meet mechanical properties, without compromising quality and aesthetics. They do so without taking it to the extreme. They avoid creating an electrode
Welders now have a broad choice of electrodes designed for a variety of specific applications and industries, expanding the range of use, as well as overall quality and productivity. With improved operating characteristics and performance, these enhanced, highly efficient products are truly not our fathers’ flux cored electrodes. ◆
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Welding Resources for When You’re on the Go A search found plenty of free, easy-to-use, welding-related apps
BY HOWARD WOODWARD, MARY RUTH JOHNSEN, CARLOS GUZMAN, AND KRISTIN CAMPBELL
S
ure you can play games, work on a crossword puzzle, count calories, track your exercise regimen, and manipulate photos through any number of applications (apps) on your smart phone or tablet, but can your work benefit from the use of an app? It so happens that when you Google “welding apps,” you’ll discover a lengthy list for download to both Android and Apple smart phones and tablets. Many apps are intuitive and require no instruction to use. The topics include selecting equipment, inspection data, weld joint design, calculators, code standards, test questions, safety, etc. After the initial downloading, most apps function without an Internet connection. Apps usually begin by asking you to answer a few questions to define your situation, then open a scroll-down page with the answers or recommended settings, data chart, a table, equipment settings, required filler metals, etc. The apps offered by equipment and consumables manufacturers make it easy to compare the features and capabilities
of each of their offerings to clarify the important features to consider when shopping for these items. While many full-featured apps can be downloaded free of charge, some free apps are “Lite” versions that offer basic versatility to promote purchasing the “Full” version. Although some are limited in scope, a Lite version may be worth checking out since the information it offers may be all you require for your daily needs. A bonus to using mobile apps is their potential educational value. You can easily “experiment” with different metals, material thicknesses, processes, etc., to get a feel for how various processes work. Following are brief descriptions of a cross section of welding-related apps available for free downloading onto your smart phone or tablet. To search for these and other apps, just Google your topic or the app’s name. You can often find useful apps offered on welding equipment manufacturers’ Web sites. Check for new apps online occasionally, since they are added and updated daily.
1
Help for Selecting Power Generators.
This well-organized, easy-to-use app details Multiquip’s line of MQ Power WhisperWatt Super-Silent portable 5- to 450-hp generators designed to provide power under the harsh conditions at construction sites, entertainment venues, and disaster-recovery operations. The data are intuitively organized into compact files that permit uncluttered viewing of images and text on the small smartphone screens. Opening the app for the first time (Fig. 1) reveals a “Begin” button. Once tapped, it opens a screen that scrolls through 19 sizes of generators from 5 to 450 hp. Tapping the hp size of your choice opens a screen showing the several voltage ranges available for that size generator. Tapping a voltage range opens a page illustrating the recommended generator including its prime, standby, and voltage dip ratings. Tapping that page opens a
HOWARD WOODWARD (
[email protected]) is associate editor, MARY RUTH JOHNSEN (
[email protected]) is editor, and KRISTIN CAMPBELL (
[email protected]) is associate editor of the Welding Journal. CARLOS GUZMAN (
[email protected]) is editor of the Welding Journal en Español. 36
MAY 2013
Fig. 1 — Multiquip MQ Power Genera- tor Selector app as it appears on a smart phone.
menu listing pages detailing this generator’s specifications, dimensions, weights, engine specifications, trailer requirements, various options, plus a “Documents” file that opens to offer a number of PDF spec sheets and equipment brochures. A “Contact Us” button also appears on this page that when touched opens into a short form for writing an email message to the company.
2
Calculator Fine Tunes Your Welding Machine.
This calculator (Fig. 2) from Miller Electric Mfg. Co. is designed to help you tune your welding machine for optimal results based on your answers to a few
simple questions. After entering the metal information, the calculator will display the suggested wire size, wire feed speed and shielding gas settings, and voltage and current ranges. The opening page of this app displays a screen with four large icons labeled MIG (Solid Wire), MIG (Flux Cored), STICK, and TIG, plus a row of icons for accessing YouTube, Facebook, Twitter, and e-mail contact. For example, touching the MIG (Solid Wire) icon asks you the following two questions: 1. What material are you welding? (The choices are aluminum, stainless steel, and steel); and 2. How thick is 1 the material? (The choices range from ⁄ 8 1 to ⁄ 2 in. and up.) After entering these data, touch the “Get Settings” icon to view the recommended settings and other suggestions.
3
This app from 3M is designed for industrial hygienists and safety professionals who want rapid access to respirator guide information on their smart phones. The guide identifies facepieces, cartridges, and filters including an overview of qualitative fit-testing protocols and how to conduct the fit tests for a variety of specific contaminants found in industrial environments — Fig. 3A. After the user enters the chemical name, the app presents the company’s recommendations for the appropriate type of respiratory protection for that contaminant — Fig. 3B. Included is a frequently asked questions section and product catalog. Currently available for
A
Fig. 2 — The opening s creen of Miller Electric’s weld s etting calculator.
App Provides Fast Access to Respirator Info.
B
Fig. 3A and B — The 3M app offers a guide for protecting yourself against a variety of specific indus- trial airborne contaminants.
the iPhone, this app is expected to be available in December for Androidbased smart phones.
4
Pipefitter’s Reference App.
The opening screen details (Welded) Flange Dimensions. The tabs shown are Series 150, 300, 600, 900, and 1500; Valve 1 Dimensions; ⁄ 2 Dimensions; Reducing Tees; Pipe Schedules; and Abbreviations. Touching the “Series Flange 600” tab, for example, opens a table cross-referencing Pipe Size, Wrench Size, Stud Length, Raised Face, Gasket, and Ring Joint. Touching the Pipe Schedules tab opens a table displaying Nominal Wall Thicknesses for commonly used pipes.
5
Ultrasound Calc Lite Version App.
The first page presents three tabs labeled Transducers Physical Principles, Characteristics of Ultrasonic Waves, and Fundamental Principles of Ultrasonic Wave Propagation. Touching the Fundamental Principles tab reveals three tabs for Snell’s Law Calculations, Geometry Calculations, and Basic Skip Calculations. Touching the Basic Skip tab, for example, opens a screen for the user to enter wall thickness of the piece, refracted angle in the piece, half skip, full skip, and sound path numbers to complete the calculations.
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6
Phased Array Wizard Lite.
This app offers a wealth of information beginning with five opening tabs labeled PA Transducer Parameters, Wedge Calculation, PA Inspection, Acoustic Parameters of Materials, and Glossary. Touching the Parameters of Materials tab opens tabs for selecting aluminum, brass, copper, glycerine, lead, nickel, Plexiglass, Rexolite, steel, titanium, and water. Touching a material’s tab presents its long and transverse velocities, density in g/cc, and the longitudinal and trans verse acoustic impedances.
7
ISMT Tube Calc Limited.
This basic app is an easy-to-use Tube/Bar Weight Calculator. The opening screen prompts you to enter two values from the menu of Diameter, Thickness, Inside Diameter, and Length. As an example, typing in 6-in. diameter, 0.25-in. wall, and 25-ft length displays a table listing the parameters with a total weight of 384.9 lb.
8
Check Live Welding Machine Status Reports.
The Lincoln Electric Co.’s CheckPoint™ app allows viewing live welding machine status and production reports on mobile devices — Fig. 4.
Once installed, it has a login process for the program’s users to input their existing username/password. Features include the following: a list of companies to see machine data and production details; an overview of all machines and live status indicator; dashboard widgets for each machine, an alerts history, and lists of all events/alarms generated by the machine; downloadable documents and manuals; historical trending data for potential issues while walking the shop floor; and scanning barcodes using the phone’s camera for operator ID, part serial number, and consumable lot code. The app works for Blackberry (supports OS 5.0 and greater), iPhone (supports iOS 4.3 and greater), and Android phones.
9
Discover Wheel Speed Calculations.
The Norton Abrasives grinding app offers calculators for a coolant, dressing parameter, and wheel speed. The portable grinder product selector has areas for selecting the application, material, and primary/secondary attributes. Postal codes can be entered in the distributor locator lookup. The application inquiry features options for general information, machine tool, abrasive and dressing products, work material, and operational factors. Also provided are sections for logging into an abrasive connection site and contacting the company — Fig. 5. The app is available for iOS and Andr oid operating systems on mob ile devices.
Fig. 5 — Among Norton’s app offerings are calculators for a coolant, dressing parameter, and wheel speed.
10 Fig. 4 — Lincoln’s app enables explo- ration of a robot’s in limit, out of limit, and total welds.
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Discuss Welding and Metalworking Topics.
The ShopFloorTalk (SFT) app by End of Time Studios, LLC, provides a discussion forum for users interested in weld-
Fig. 6 — The ShopFloorTalk app has many forums, including one for welding processes.
ing, fabrication, machining, and general metalwork of all kinds — Fig. 6. The welding and metalworking forums offer the following topics: fabrication; equipment, suppliers, and original equipment manufacturers; welding processes; machining; metallurgy and materials; shop safety; shop; general welding info rmation; SFT information, workshops, and support; and b usiness. These sections have numerous posts, including photos of finished, painted art works; questions and answers from what welding process to use and when; tips for getting customers; and much more. There is also a members only forum. To post messages, users must register; provide profile information; and then activate their account. The app is compatible with the iPhone, iPod touch, and iPad; requires iOS 4.3 or later; is optimized for the iPhone 5; and Android (1.5 and up).
11
Select Job Materials for the Metalworking Industry.
The Tool Steel Selection Guide app by Lindquist Steels, Inc., is for tool makers and designers, engineers, and users interested in selecting and utilizing the company’s materials — Fig. 7. The build tab asks users the following question: What would you like to create
type in the thickness of the weld and it calculates the most common discontinuities including cracks, incomplete penetration, internal porosity, slag, undercutting, and concave root. It also calculates the sizes of image quality indicators (IQIs) needed. The IQI area has three sections for welds with no reinforcement and welds reinforced on one or both sides. Lance Henderson, the app’s developer, offers four o ther free apps: X-Ray Timer, Density, Ug Calc, and Barricade. Information is available at www.lanc e hen derson.co m as well as through the Apple app store.
Fig. 7 — Lindquist’s app acts as a ref- erence guide to work around tool pro- duction needs. today? An A–Z list has keywords from aluminum extrusion tooling to zinc die cast dies. By clicking on each term, suggested materials from low to high production value are found. Another area for materials can also be explored. In addition, users can obtain a quote/order, discover the company’s history, and find contact details. The app is compatible with the iPhone (3GS, 4, 4S, and 5); iPod touch (3rd, 4th, and 5th generations); and iPad.
12
Fig. 8 — Users plug a few pieces of in- formation into ESAB’s Welding Parame- ters Set-Up Guide and receive data that can be used to set up their welding ma- chine and adjust it to their specific ap- plication requirements.
13
Help with Interpreting Industrial Radiographs.
The Code 313 app (Fig. 9) is a calculator for interpreting industrial radiographs in accordance with ASME B31.3, Pro cess Piping . The app’s developer warns that the app does not replace the code itself, nor the need to have the code, but is a calculator to help you verify your own interpretation. To use the app, you
Identify the Welding Parameters for Your Next Job.
The Welding Parameters Set-Up Guide app from ESAB Welding & Cutting Products provides users with the welding parameters they need for a particular job, including wire feed speed, voltage, current, and inductance — Fig. 8. The first step is to select solid wire or flux-cored wire at the bottom of the screen. If the user selects solid wire, it will then ask what type of material, thickness, wire type, and gas type will be used (each of these are selected through scroll-down listings). The guide then lists the recommended settings. If flux-cored wire is chosen, the guide asks users to select material thickness and wire type before listing the settings. The app is compatible with iPhone, iPod touch, iPad, Blackberry, and Android devices.
Fig. 9 — Code 313 helps with interpreta- tions of industrial radiographs.
14
Calculates Weld Costs.
The Welding Pro app by Certilas Nederland BV, a maker of welding consumables in The Netherlands, provides weld cost calculations for fillet welds, and single-V, double-V, and double bevel butt joints — Fig. 10. It allows users to quickly compare labor, gas, and filler metal costs. Users pick the type of weld they will be making (fillet or groove), then select welding process, amperage, duty cycle, type of electrode, etc. Users also designate whether they are welding steel, stainless steel, or aluminum. While the app is set up for metric units and euros, a tap on the settings button allows users to switch to Imperial Standard units and dollars. It also displays the most recent calculations and users can designate certain calculations as “favorites” that can be stored and retrieved later. Requires Android 2.2 and up or iOS 4.3 or later.
Fig. 10 — Welding Pro allows users to quickly compare labor, gas, and filler metal costs.
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15
Troubleshoot Resistance Spot Welding.
The Resistance Spot Welding Troubleshooting App by Miyachi Unitek offers a quick guide to solving common problems found in the RSW process — Fig. 11. The sharp interface is simple to use and is well laid out. There are four main menus or tabs: Instructions, Guidelines, Troubleshoot, Troubleshoot, and Info. By tapping on the Troubleshoot tab, the user is able to choose from a variety of possibl e symptoms or problems: overheating of weldment, discoloration, weak weld, insufficient nugget, metal expulsion, sparking, inconsistent welds, electrode damage, and electrode sticking. Solutions are presented in four categories: material, electrode, weldhead, and power supply. Additio Addi tionall nally, y, the user u ser sel selects ects one of four priority levels, according to the most likely cause of the problem. For example, if overheating of weldment is the most likely cause of the problem, select Priority 1 and the app reveals that it may be caused by excess time. This suggestion appears in the Power Supply Related field. Tap on this field and a possible solution is shown (in this case, “Decrease weld time in steps of 5–10% 5 –10% ...”). Althoug Alth ough h the th e soluti so lutions ons pres presente ented d in this app are brief and may not cover all vari abl ables, es, it’ it’ss a goo d star startin tingg poi point nt for professionals and a worthwhile tool for educational purposes. This Miyachi Unitek app is available for iPad, iPhone, and Android.
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Growing Catalog of Welding Calculat C alculators. ors.
AxonCal Axo nCalcc has deve develop loped ed a seri series es of app calculators for welding, mechanical, and materials engineering — Fig. 12. The Welding AxonCalc app includes the following calculators: Carbon Equivalent and Composition Parameter (PCM), PreHeat and Interpass Temperature, Maximum Hi-Lo at Internal Diameter for Pipes, Hi-Lo at Internal Diameter at Specific Location for Pipes, and Arc Welding Heat Input Calculator. Designed with a legible and attractive interface, the AxonCalc calculators are easy to navigate and use. For example, the Arc Welding Heat Input calculator
17
Look Up Cut Charts and Error Codes.
The Thermal Dynamics Cut Chart is designed for the company’s Ultra-Cut series of high-precision automated plasma systems — Fig. 13. The app offers a cut chart for Ultra-Cut with DPC3000 and error codes for all UltraCut as well as Auto-Cut systems — Fig. 13. Under the best cut division, users can select the appropriate material type, including aluminum, mild steel, and stainless, along with the thickness (gauge, in., or mm). Corresponding options will appear, and some materials have multiple options. Entering error codes will enable viewing troubleshooting troubleshoot ing suggestions. Also, there are sections sec tions linking to the Victor Technologies Technologies Cutting & Welding YouTube channel and locating technical
Fig. 12 — AxonCalc welding calculators have a clean interface that is easy to use.
Fig. 11 — By tapping on the Troubleshoot Troubleshoot tab, the user is able to choose from a va- riety of possible symptoms or problems.
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has fields for inputting current, voltage, and travel speed alone or travel speed and time. The results, which can be copied onto the device’s memory or e-mailed directly from the application, are displayed in both kJ/in. and kJ/mm. This free app contains demos of all the welding calculators, which lets the user try them before committing to buy, and the company has plans to expand the collection regularly. Additional app demos can be found on the company’s Web site: www.axoncalc.com.
Fig. 13 — The Thermal Dynamics Cut Chart features areas for looking up a cut chart along with error codes.
support in Australia, Canada, China, Europe, South East Asia, and the United States. The app is compatible with the iPhone (3GS, 4, 4S, and 5); iPod touch (3rd, 4th, and 5th generations); and iPad. It requires iOS 5.0 or later. ◆
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Improving Surfacing Performance with GMAW The synchronized polarity gas metal arc welding technique shows promise for applying coatings where the risk of perforating the base material is a critical factor
new technique for the production of a coating using the gas metal arc weldi ng (GMAW) ( GMAW) pro cess is prepr esented in this article. The technique fulfills two important requisites for this type of operation: low dilution and high productivity. In order to carry out a welding task, some characteristics associated with the process and its respective procedure need to be sought in order to fulfill the technical requirements specified. In the
A
BY JAIR CARLOS DUTRA, EDUARDO BIDESE PUHL, NELSO GAUZE BONACORSO BONACORSO,, AND REGIS HENRIQUE GONCALV GONC ALVES ES E SILVA JAIR CARLOS DUTRA, EDUARDO EDUARDO BIDESE PUHL (
[email protected]) , , and (
[email protected]) REGISS HENRI REGI HENRIQUE QUE GONC GONCAL ALVES VES E SIL SILVA VA are with Federal University of Santa Catarina — UFSC Mechanical Engineering Department, Florianópolis, Brazil. NELSO GAUZE BONACOR BONACORSO SO is with Federa Federal l Institute of Education, Science and Technolog Technologyy of Santa Catarina, Metal Mechanics Department, Florianópolis, Brazil.
case of coating applications, beside the absence of defects in the weld beads and in their overlaps, penetration and dilution must be minimized in order to guarantee the intended quality. Despite the fact that this study encompasses the parameters mentioned above, the intention herein is not to apply them as comparison parameters. What is of fundamental relevance is the possibility of using the GMA GMAW W process with direct current electrode negative (DCEN)
polarity, a welding condition traditionally considered inappropriate due to the instability of the arc and a weld bead geometry that is completely unsuitable. The instability is currently resolved through the use of a specific gas composition within a certain range of electrical current. The geometry of the deposit is then solved by the synchronized polarity gas metal arc welding (SP-GMAW) (SP-GMAW) technique proposed here, which consists of the synchronization of the welding power
Table 1 — Weld Beads Obtained in the Flat Position on an ABNT 1020 Carbon Steel Plate (AWS ER70S-6) Wire with Diameter of 1.2 mm, Mixture of Ar and O 2, Welding Speed of 348 mm/min, and Current of 250 A
Test
Polarity
1
DCEP
7.2
2
DCEN
11.7
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Feed Speed (m/min)
Weld bead appearance
Cross section
A
source output polarity with the torch position in relation to its oscillatory motion. Welders discard the use of DCEN with solid electrodes as an option because of the allegation that the electric arc is unstable and produces a globular metal transfer with a significant amount of spatter. The geometry of the bead deposited is unsuitable due to the low wettability, which results in an almost circular cro ss section that can lead to discontinuities due to incomplete fusion at the joints with adjacent weld beads. However, for coating operations, other characteristics, such as low levels of penetration and dilution, are required. Recent studies on the GMAW process with DCEN polarity have verified, via high-speed digital filming, that the type of shielding gas has a significant influence on metal transfer behavior and weld bead geometry (Ref. 1). Good results were obtained with an appropriate composition of argon and oxygen. The metal
B
transfer was globular at 150 A and axial spray at 250 A without drop repulsion, in contrast to descriptions in the classical literature. The bead profile has low wettability (inappropriate format) but acceptable penetration, as shown in Test 2 (Table 1). The use of DCEN in the GMAW process is a real possibility for tubular electrodes with a slag-forming flux, but in the case of solid electrodes, it is limited only to the context of alternati ng current. In this latter case, the frequency of the polarity switching is associated with the frequency of the drop transfer and, coincidently, is close to the frequency (50 or 60 Hz) the electricity-generating companies provide. One sought-after property is the greatest amount of molten electrode material for a certain arc power. This provides the process with particular characteristics in order to achieve specific objectives. For example, researchers at Fronius, a manufacturer
Fig. 1 — Electric arc showing the behavior of the GMAW arc in the following polarities: A — DCEP; B — DCEN.
of welding power sources including CMT Advanced (cold metal transfer), revealed the gap-bridging capacity of the molten material of the joint faces in root passes, which they designated as the bridgeability (Ref. 2). Anot her important characteristic of DCEN that is favorable for the application of the coating is the high wire-melting rate for a certain current in comparison with direct current electrode positive (DCEP) polarity. This distinctive melting rate, also represented by the differentiated wire feed speed in Table 1, can be explained by the behavior of the electric arc. In DCEN polarity, the electric arc does not anchor only at the end of the electrode as in DCEP (Fig. 1A), but instead, widely embraces the electrode (Fig. 1B), seeking points where the electron emission is favorable for the presence of oxides. This characteristic leads to a greater parcel of the arc energy being transferred to the electrode, enhancing its relative melting at the expense of the melting of the workpiece (Refs. 3, 4). The objectiv e of the new SP-GMAW technique is to minimally affect the base material and provide high productivity with a reduced amount of defects. A potential application for this technology is the repair of in-service pipelines that undergo a reduction in wall thickness due to corrosion. In this application, more than the need for low dilution, the obtainment of highly reduced penetration and the possibility of carrying out the repair in a short time are of fundamental importance. However, this type of coating is still carried out with coated electrodes, where the result is fundamentally dependent on the ability of the welder and the execution time is long.
The SP-GMAW Technique The developed technique was given the denomination of synchronized polarity (SP) due to the characteristic of the change in polarity during the GMAW process in synchrony with the torch po-
Fig. 2 — Functioning strategy for the SP-GMAW technique (Ref. 5).
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sition in the trajectory of the weld bead execution. Robots or manipulators automatically execute the weld beads, and combine the qualities of both polarities in the GMAW process. The negative polarity is used in the center of the bead tra jectory in order to ob tain greater melting rate and welding speed, and lower dilution and penetration. The positive polarity is used only at the ends of the oscillation trajectory to prepare the weld bead for adequate overlapping with another that will be deposited alongside it, thus avoiding fusion defects. In the qualification of the procedure the values for Yt and Yp of the transversal Y axis (Fig. 2) need to be appropriately considered. The aim is to obtain weld beads with the geometric characteristics shown in Fig. 3. The task of synchronizing the polarities with the oscillation trajectory of the gun is carried out via a digital synchronization signal generated in the manipulator controller and recognized at the welding source, as shown in Fig. 4. Thus, the welding equipment involved — manipulator and power source — must be based on digital technology with the possibility for the programming and parametrization of electrical signals. For a certain weld speed and oscillation amplitude/frequency, configured at the programming interface of the manipulator, the position of the transversal Y axis is constantly compared with the transition amplitudes –Yt and +Yt, as shown in Fig. 5. When the position of the Y axis surpasses one of the transition amplitudes, –Yt or +Yt, the synchronization signal shifts to logic level 1, which, in turn, commands the welding power source to impose the I+ current on the electric arc through the application of DCEP polarity. If the position of the Y axis is between these transition amplitudes, the synchronization signal switches to logic level 0, and commands the welding source to impose the I– current on the electric arc through the application of DCEN polarity. The value in the current module imposed by the source, I+ or I–, must be in agreement with the respective wire feed speeds established, since the melting rates are different for each polarity when the current values are the same. If the reaction dynamics of the wire feeder used is low, then the same wire speed can be applied for the two polarities. In this case, the process equilibrium is achieved through the adjustment of the intensity of each current, I+ and I–.
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Fig. 3 — Characteristics expected of the weld bead cross section.
Fig. 4 — Functioning diagram of the SP-GMAW technique.
Fig. 5 — Synchronization logic of the SP-GMAW technique.
4
5
3
Fig. 6 — The equipment setup for demonstrating the validity of the synchronized polarity gas metal arc welding technique.
Setting up the System The tests performed to demonstrate the validity o f the SP-GMAW technique simulated the recovery of the thickness of a worn low-carbon steel workpiece through the addition of 1.2-mm-diameter ER70S-6 wire in the flat welding position. The equipment used for the tests consisted of a microprocessor-controlled welding power source (Model IMC In versal 450) with its respective wire feeder, a Cartesian XY manipulator (Model SPS Tartílope V2F), and a portable data-acquisition system (Model
IMC SAP-V4.01) (Ref. 6) — Fig. 6. The welding power source was programmed to operate in GMAW mode with the current imposed instead of the voltage. Thi s produces better dynamics for the alternating polarities. The shielding gas used was a mixture of 98%Ar + 2%O 2 with a flow rate of 13 L/min that was recommended by researchers (Ref. 1). Table 2 shows the parameters of the movement and of the welding used to produce the weld bead shown in Fig. 6. This deposit was obtained using the push technique with an angle of 10 deg. The
Table 2 — Values for the Parameters Applied in SP-GMAW Parameters Weld speed (mm/min) Oscillation frequency (Hz) Oscillation amplitude, 2.Yp (mm) Polarity amplitude DCEN, 2.Yt (mm) Wire feed speed (mm/min) Electric current in DCEP polarity, I+ (A) Electric current in DCEN polarity, I– (A)
Values 436 1.5 12.0 8.0 11.5 270 250
application time for DCEN was twice that of DCEP. Despite the occurrence of spattering (Fig. 7A), a regular weld bead with an average width of 15.4 mm and maximum height of 3.1 mm was obtained. The maximum penetration (Fig. 7B) was only 0.5 mm and the dilution was approximately 13%. The oscillograms of the voltage and current of the electric arc, as well as the wire feed speed (Fig. 8), were captured during the production of this particular weld bead. The temporal behavior of the electric arc voltage indicates that short circuiting does not occur. The spattering observed can be explained by the projection of drops of the electric arc to the outside of the weld pool due to the inversion of the direction of the weld torch movement. This inversion of movement occurs during the application of DCEP polarity and at both lateral ends of the trajectory of the torch oscillation. It can also be observed that in DCEP polarity, when the I+ current reaches its reference value (270 A), the voltage gradually reduces. This signals the decrease in the length of the electric arc due to a lack of wire consumption. When the poWELDING JOURNAL
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larity is switched to DCEN, the length of the electric arc is reestablished, indicated by an increase in its voltage. In this case, the electric arc is stable, but operating close to the lim it of instability . To make the process more stable, the DCEP current of the DCEP polarity (I+) must be corrected, that is, increased until an appropriate wire feed speed is reached. Also , other factors, such as nonlinear electromagnetic force variations of the arc, influence the transitory characteristics that can contribute to the previously mentioned voltage variation. Another unstable and undesirable situation is when one or both currents, I+ and I–, are higher than the currents suitable for a certain wire feed speed. In this case, the arc length increases excessively and melts the contact tip of the torch. The same set of parameters used in the previous test (Table 2) was applied to make the weld shown in Fig. 9. In order to ensure the overlap of the weld beads only in the region of DCEP polarity, the following criteria were applied: distance between the longitudinal axes of two ad jacent weld beads equal to the value of the oscillation amplitude, that is, 12.0 mm. The layer obtained (Fig. 9A) has good superficial appearance with a maximum height of 3.1 mm and a maximum undulation of less than 0.3 mm. Its cross section (Fig. 9B) reveals a shallow penetration and the absence of weld defects.
A
B
Fig. 7 — A — The weld bead obtained with the SP-GMAW technique; B — cross section of the weld bead.
Conclusions The SP-GMAW technique offers a real possibility for the application of coatings where the risk of perforating the base material is a critical factor. This is due to the achieved appropriate geometric characteristics for the task of coating surfaces, such as shallow penetration, a surface with almost no undulation, and good dimensional ratio (width/height) of the weld beads. The coating criteria adopted in the SPGMAW technique for the overlap of ad jacent beads produced good results in the flat welding position. There is no increase in the height of the weld layer due to this overlap. The maximum undulation generated was lower than 0.3 mm. The weld layers produced do not present discontinuities and have an excellent visual aspect with the presence of very little spatter. The use of DCEN polarity during times that extend beyond the period of drop transfer should not be discarded. If used with certain wires in conjunction 46
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Fig. 8 — Behavior of the main variables of the SP-GMAW technique.
A
B
Fig. 9 — A — The coating obtained with the SP-GMAW process; B — a cross section of the coating.
with adequate gas mixtures and a suitable current range, it can represent a good alternative for specific cases of welding. However, this technology requires pieces of equipment that communicate with each other and considerable dedication in the qualification of the set of variables and parameters for the GMAW process as well as for automatic torch displacement. ◆
Acknowledgments
The authors are grateful to the team at Labsolda/UFSC whose efforts made this multidisciplinary task possible and to the Brazilian government agency CNPq and the company Tractebel Energia for financial support.
References
1. Souza, D., Rezende, A. A., and Scotti, A. 2009. A qualitative model to explain the polarity influence on the fu-
sion rate in the MIG/MAG process. Re vis ta Soldagem & Ins peção 14(3): 192–198. 2. Pickin, C. G., Willams, S. W., and Lunt, M. 2011. Characterization of the cold metal transfer (CMT) process and its application for low dilution cladding. Journal o f Materi als P rocessin g Technol ogy 211(3): 496–502. 3. Ueyama, T., Tong, H., Harada, S., and Passmore, R. 2005. AC pulsed GMAW improves sheet metal joining. Welding Journal 84(2): 40–46. 4. Cirino, L. M. 2009. Study on the effects of polarity in direct and alternate current TIG and MIG/MAG welding processes. Master’s thesis. PosMec/ UFSC, Florianópolis, Brazil. 5. Puhl, E. B. 2011. Development of MIG/MAG welding technologies for productivity and quality enhancement by means of negative polarity. Master’s thesis. PosMec/UFSC, Florianópolis, Brazil. 6. Labsolda: Operation Manuals. Available at www.labsolda. ufsc.br/ projetos/manuais/manuais.php . Instituto de Mecatrônica. Accessed June 14, 2012.
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Exploring the Forces that Shape Droplets during Gas Metal Arc Welding A model is presented to better understand the physics of drop formation
BY ALEXEI YELISTRATOV ALEXEI YELISTRATOV
(
[email protected]) is a research associate.
ew, effective variations of the gas metal arc welding (GMAW) process are emerging. Some of these methods, such as variations of controllable drop transfer, pulse welding methods, etc., are based on controlling the drop transfer mode. These methods improve the performance of GMAW for thin-metal, root-pass welding by applying controllable pulses of current and thereby changing conditions for forming
N
A
the liquid metal drops. Because of this, the physical processes causing the drop transfer mode changes are of key importance for further development of GMAW technologies, which attracts the attention of multiple researchers (Refs. 1, 2, 4). The application of gas mixtures of argon (Ar) with carbon dioxide (CO 2) or oxygen (O 2) expands the arc and increases the anode spot size while decreasing droplet diameter (Refs. 1, 3). By
gradually increasing the welding current (A), the transition to spray mode occurs sharply, within a transition current range of 50–80 A — Fig. 1A. For GMAW in Ar-rich gas mixtures, the transition current density is 120–190 A/mm2 (Ref. 3). When Ar + oxidizing gas (1 to 5% O2 or 5 to 8% CO 2) mixtures are used with a steel wire, the stability and the range of technological modes demonstrate marked improvement when com-
B
B
1 16 -in. (1.6-mm) welding wire Fig. 1A — Effect of welding current and shielding gas on droplet frequency for GMAW with ⁄ 1 16 -in. (1.6-mm) welding (Ref. 3); B — effect of welding current (current density) on drop transfer mode for GMAW with ⁄ wire ER70S-3 in shielding gas mixture Ar + 5% O2 (Ref. 7).
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A
B
Fig. 2 — Directions of the electromagnetic force acted on molten droplet at globular and spray modes. Current-carrying areas at the bottom of the electrode correspond to areas of the arc anode spot.
pared to an arc with single-gas shielding. Accordi ng to multiple observations, anode spot expands with increased current (this provides its constant current density), and in spray mode it envelops the droplet and the tapered bottom part of the wire — Fig. 1B. In pulsed GMAW of aluminum, controllable droplet and spray modes can be achieved (Ref. 6). In this study, an ab rupt transition from pulsed-globular to pulsed-spray transfer is mentioned (see conclusion #3 and Ref. 6).
The Metal Transfer Forces The physical forces in the electric arc zone that are mainly responsible for droplet transfer at the transition current range are electrodynamic (pinch-effect) and surface tension (Ref. 8). The electrodynamic force acts as a squeezing force, constricting the conductor when the current-carrying cross sec-
Fig. 3 — Experimental setup for hydraulic modeling of droplet formation. 1 — Upper vessel; 2 — thermocouple; 3 — extension; 4 — lower vessel with water; 5 — electrical heater. A — Position of lower vessel (with boiling water) that corresponds to highest vapor concentration at the droplet forming zone; B — in itial position of the lower vessel.
tions are the same. There is, however, another important feature of this constricting force: When current-carrying cross sections are different, an axial electrodynamic force is created. This force directs from lesser cross section to the larger one. Welding current undergoes change in the current-carrying cross sections during its path from the welding wire to the anode spot — Fig. 2. This is similar to “conduction angle” (Refs. 3, 5). Electrodynamic forces can be directed up (Fig. 2) when the anode spot size is smaller than the diameter of the electrode at low-current globular transfer in GMAW. In this case, the electrodynamic force and surface tension force acted jointly to support the droplet on the tip of the electrode (wire). This force can be directed down when the anode spot area becomes larger than the cross section of the electrode at high-current, spray transfer with GMAW in Ar+CO 2 /O 2 mixtures (Refs. 4, 5).
During GMAW with Ar+O 2 /CO 2 mixtures, the anode spot is stabilized at the bottom of the metal droplet (Refs. 1, 3), and heat energy is transferred from the arc to the electrode through the bulk of the droplet. With a further increase in the welding current, the anode spot size becomes larger than the external surface of the droplet, eventually expands to include the cylindrical surface of the electrode, and melting occurs radially, creating a taper (Refs. 1, 4). The temperature inside the anode spot is approximately the metal boiling point (Ref. 4), resulting in intensive evaporation of the metal. The increased droplet temperature (Ref. 7) and increased metal vapor concentration in the arc zone could lead to increased electrical conductivity in the arc zone and result in sharpening the electrode tip (Ref. 8). In summary, the electrodynamic force changes direction and becomes directed downward during GMAW in Ar-rich WELDING JOURNAL
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Fig. 4 — Water modeling. A — Droplet diameter vs. vapor concentration; B — droplet frequency vs. vapor concentration.
A
shielding gases when the welding current exceeds the transition current. There is no information about variation of surface tension force for GMAW at the transition current range. The purpose of the following experiments was to study the relationship between welding current and anode spot size, and the influence of surface tension on droplet formation in conditions similar to ones at the tip of the molten electrode within transition current range.
Experimental Procedures The experiments were performed using direct current, electrode positive (DCEP) GMAW head with constant voltage (CV) power supply. Used throughout these experiments were mild steel plates 7.87 × 2.36 × 0.39 in. (200 × 60 × 10 1 mm), AWS ER 70S-3 welding wire ⁄ 16 in. (1.6 mm) diameter, and various shielding gases (100% CO 2; 20% CO2 + 80% Ar; 10% CO2 + 90% Ar; 5% CO 2 + 95% Ar). The shielding gas mixture flow rate was 42 ft3 /h (20 L/min). The welding current and voltage were recorded. Highspeed photography (4000 f/s) was used to capture droplet measurements. The geometrical sizes of the droplets and droplet frequency were measured from the image on the screen. The anode spot size was measured from the image on the screen by measuring the size of the bright spot on the outer side of the droplet. This indirect method is acceptable for evaluation of low-scale objects in the electric arc zone (Ref. 4). The variations in measured droplet size were ±5%, and in anode spot size were ±10–15%.
Computational Model of the Surface Tension Force Significant measurement errors can occur when studying the arc zone conditions using direct methods such as probes, thermocouples, etc. Practical calculation of the surface tension force for GMAW conditions is impossible because there are no reliable supporting data. Therefore, in this work, a qualitative method was used to study the influence of 50
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B
the surface tension force — physical modeling of liquid droplet formation. The physical model is not a model for GMAW, nor a model for GMAW metal drop transfer; instead, it is just the distinguishing group of physical interactions where the surface tension force predominates. This model allows investigating only the effect of surface tension and eliminates the influence of forces related to electrical current. Selection of the modeling liquid and analog model dimensions were made in accordance with theory of similarity. Although calculations have been made for a number of liquids, water was selected as the most suitable for these experiments. Water and low-melting alloys are used to study the flow of liquid steel in molds and feed head systems in the metallurgical industry; also, a water/alcohol model was used (Ref. 9) to study droplet/ molten pool interaction in GMAW.
This computation objective is to study the surface tension variations in conditions similar to real welding, using water as a modeling liquid. Conditions that are similar in forming both, real arc welding metal drop, and water drop are: • Drop temperature is close to the boiling point. • The growing drop is surrounded by the drop’s liquid vapors. Since both liquids, molten steel and water, have negative coefficients of surface tension, they react similarly to variations in the temperature of the dropforming zone: The higher the temperature (closer to boiling), the lower the surface tension. The surface tension is caused by cohesion forces inside the liquid and acts on any liquid surfaces in accordance with the liquid’s physical properties. To investigate the influence of surface tension on water droplet transfer, the
Fig. 5 — Droplet external surface and anode spot area at GMAW. A — 100% CO2 ; B — 95%Ar + 5%CO 2.O2.
A
tration of vapor around the drop, not the surrounding temperature. For evaluation of the water vapor concentration, additional experiments were conducted to measure the vapor concentration (humidity) with a portable thermohygrometer Testo 605-H1, which was used in parallel with measuring the temperature in the drop-forming zone. According to results, the hygrometer cannot provide accurate readings due to the restricted space for measurements (less than 1 cm 3), but for larger spaces the readings from the hygrometer and thermocouple correlated to 5 to 7 % at a fixed distance for the thermocouple of a few millimeters from the drop surface. Also, it was found the correlation between vapor concentration and temperature improved with approaching the boiling point. Because of that, temperature readings from the thermocouple were accepted to evaluate the vapor concentration present in the small drop-forming zone.
B
Effects of Water Vapor
model (Fig. 3) included an upper vessel with water, an extension wherein the droplets form, and a lower vessel with boiling water and an electric heater. The experimental setup included a thermocouple, which was fixed closely to the outer end of the extension to measure the temperature at the droplet-forming zone. The extension used a special porous insert to provide the laminar flow of the water during droplet formation. The concentration of the water vapor was determined indirectly, by temperature in the droplet-forming zone, i.e., the temperature of boiling water (100°C) corresponds to 100% vapor concentration near the surface. The formation of the water droplets was observed using a 25× microscope. The water model for studying surface tension effect in conditions compatible to that at transition current range in GMAW has several features:
1. In the model, the drop is formed by water flowing from top, while in real welding, wire melts from its bottom by action of an electric arc. Current-caused forces (pinch-effect, plasma jets, etc.) make the real drop formation more complicated, but the resulting effect of that is variations in the drop temperature and surface tension since they are the main forces that support the drop. 2. When the heater moved up, radiated heat could heat the upper vessel with water. But with the thermocouple installed near the drop, the total surrounding temperature was registered without relation to ways of heating. To minimize that effect, the time for the experiment was limited and each set of experiments was repeated 8 to 10 times. 3. The method for evaluating the vapor concentration by temperature in the drop-forming zone cannot be accurate but what is important is the concen-
In this set of experiments, the lower vessel with boiling water was moved up, closer to the droplet-forming zone. This allowed the temperature and vapor concentration in the droplet-forming zone to reach 100°C (100% vapor). The temperature in the droplet formation zone and the droplet transfer parameters (diameter and frequency) was determined at various vapor concentrations — Fig. 4. When the concentration of the water vapor approached 100%, sharp changes in droplet transfer parameters were observed. The drop diameter decreased from 0.177 to 0.039 in. (4.5 to 1 mm), and the drop transfer frequency increased from 40 to 90/s.
Effects of Temperature An electrical heater without the lower vessel was installed below the droplet formation extension and moved up toward the extension. When the heater was in the upper position, the size of the droplets changed. The droplets lengthened in the direction of the heater from 0.019 to
WELDING JOURNAL
51
0.047 in. (0.5 to 1.2 mm), and their diameters decreased from 0.177 to 0.149 in. (4.5 to 3.8 mm). The drop transfer frequency increased from 40–43/s to 50–53/s.
Effects of Adding Dyes and Ceramic Powders Aniline dye, which is a soluble liquid, was deposited on the surface of a droplet. In a fraction of a second, the dye had spread through the bulk of the drop. The amount of aniline dye used did not affect the diameter or the frequency of droplet transfer. Mineral oil, a nonsoluble liquid, collected at the bottom of a water droplet, and formed an independent drop. When the oil was injected into the water droplet (through a thin pipe), an independent oil sphere formed inside the water droplet. There were no noticeable changes in drop transfer. When dry ceramic powder was deposited on the surface of a water droplet, it occupied only the surface of the droplet. The powder, previously soaked in aniline dye, penetrated through the bulk of the water droplet. Powder, previously soaked in mineral oil, occupied only the surface of the droplet. When a mixture of ceramic powders preliminarily soaked in aniline dye and in mineral oil was used, penetration into the bulk of the water droplet depended on the composition of the mixture. When there was more than 30 to 50% dyed powder in the mixture, the powder began to penetrate into the bulk of the water droplet, carrying the oiled powder. There were no noticeable changes in drop transfer. From the modeling experiments, droplet transfer parameters and therefore surface tension are not dependent considerably on temperature in the droplet-forming zone until the temperature in drop forming zone approaches the boiling point. After that, the high concentration of vapor causes the droplettransfer parameters to change in a marked degree because of the drastic decrease in the surface tension force. An analysis of the high-speed photographs of the droplet transfer during GMAW in different gas mixtures (Fig. 5A, B) confirmed that change in droplet transfer mode from globular to spray (Fig. 1) occurred at the instant the anode spot on the droplet’s surface became greater than the droplet external surface (Fig. 5B). In this study, this occurred when the following parameters were 52
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used: gas mixture 95% Ar+5% CO 2, welding current 290 to 310 A, and current density 145 to 160 A/mm 2.
Model Explains Iron Powder Effect on Stability The modeling experiments developed in this study explain the change in droplet transfer mode when welding within the transition current range as follows: • As the axial electrodynamic (“pincheffect”) force is directed down and gradually increases, the anode spot size becomes larger than the cross section o f the electrode and larger than the external surface of the droplet. • At this moment, the value of the surface tension decreases drastically because formation of the droplet occurs completely inside the arc’s anode active spot, within the zone with high metal vapor concentration. Together, those two factors lead to an abrupt change in droplet size, changing the transfer mode from globular to spray. Experiments confirmed that any method that increases the metal vapor concentration in the arc zone would decrease the surface tension force of liquid metal. For example, the addition of iron powder into the flux core (or in the electrode coating), in conjunction with increased deposition rate, will probably decrease the droplet diameter through the intensive metal vaporization, thus providing a more stable process. Also, modeling experiments with ceramic powder mixtures (conditions similar to flux core and shielded metal arc welding) demonstrated that interaction between liquid metal and slag are dependent on the composition of the slag.
Some Interesting Results 1. For GMAW with Ar-rich shielding gas, increasing the anode spot size above the droplet external surface provides high metal vapor concentration in the droplet-forming zone. 2. The main cause of transfer mode change from globular to spray is the increase in the electrodynamic (“pinch effect”) force and the sharp decrease of the liquid metal surface tension force on the tip of the melted welding wire. 3. During GMAW in the globular transfer mode, the surface tension force and the electrodynamic force are two important factors in forming of the metal drops. In the spray mode, when welding current exceeds the transition value, the
influence of the surface tension force on droplet formation becomes insignificant and the electrodynamic force governs the liquid metal transition to the molten pool. 4. The surface tension force can be another effective tool for controlling the droplet transfer, arc stability, and alloying efficiency. ◆
Acknowledgment
The author would like to acknowledge Dr. A. Lesnewich for reviewing the manuscript and for his valuable suggestions.
References
1. Soderstrom, E. J., and Mendez, P. F. 2008. Metal transfer during GMAW with thin electrodes and Ar-CO 2 shielding gas mixture Welding Journal 87(5): 124-s to 133-s. 2. Hu, J., and Tsai, H. L. 2006. Effects of current on droplet generation and arc plasma in gas metal arc welding . Journal of Applied Physic s , 100, Article No. 053304. 3. Rhee, S., and Kannatey-Asibu, E. Jr. 1992. Observation of metal transfer during gas metal arc welding. Welding Journal 71(10): 381-s to 386-s. 4. Kim, Y. S., McEligot, D. M., and Eagar, T. W. 1991. Analysis of electrode heat transfer in gas metal arc welding. Welding Journal 70(1): 20-s to 31-s. 5. Jones, L. A., Eagar, T. W., and L ang, J. H. 1998. Images of a steel electrode in Ar-2%O 2 shielding gas during constant current gas metal arc welding. Welding Journal 77(4): 135-s to 141-s. 6. Subramaniam, S., White, D. R., Jones, J. E., and Lyons, D. W. 1998. Droplet transfer in pulsed gas metal arc welding of aluminum . Welding Journal 77(11): 458-s to 464-s. 7. Soderstrom, E. J., Scott, K. M., and Mendez, P. F. 2011. Calorimetric measurement of droplet temperature in GMAW. Welding Journal 90(4): 77-s to 84-s. 8. Wang, F., Hou, W., Kannatey-Asibu, E., Schultz, W., and Wang, P. 2003. Modelling and analysis of metal transfer i n gas metal arc welding. Journal of Physics D: Applied Physics 36, pp. 1143–1152. 9. Choo, R. T. C., Mukai, K., and Toguri, J. M. 1992. Marangoni interaction of a liquid droplet falling onto a liquid pool. Welding Journal 71(4): 139-s to 146-s.
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COMING EVENTS
NOTE: A DIAMOND ( ♦) DENOTES AN AWS-SPONSORED EVENT.
12th Annual Great Designs in Steel Seminar. May 1. Laurel
Manor Conf. Center, Livonia, Mich. Sponsored by The Steel Market Development Institute. www.sasft.org/en/sitecore/content/Autosteel_org/Web%20Root/Great%20Designs%20in%20Steel.aspx. ♦ JOM-17, Int’l Conf. on Joining Materials. May 5–8. Konventum
Lo Skolen, Helsingør, Denmark. Institute for the Joining of Materials (JOM) in association with the IIW. Cosponsored by AWS, TWI, Danish Welding Society, Welding Technology Institute of Australia, University of Liverpool, Cranfield University, Force Technology, and Brazilian Welding Assn. www.jominstitute.com/ side6.html. AISTech 2013, Iron and Steel Technology Conf. and Expo. May
6–9, Pittsburgh, Pa. www.aist.org/aistech/ . INTERTECH 2013, Superabrasive Materials, Principles, and Applications for the Aerospace and Defense Industries. May 6–8.
Hyatt Regency Baltimore Harbor Hotel, Baltimore, Md. Industrial Diamond Assn. www.intertechconference.com. POWER-GEN India & Central Asia, Renewable Energy World Conf. & Expo, and HydroVision® India. May 6–8. Bombay Exhi-
American Welding Society, Fabricators and Manufacturers Assn, Int’l, Society of Manufacturing Engineers, and Precision Metalforming Assn. www.aws.org/show/weldmex2013.html . Int’l Thermal Spray Conf. and Expo. May 13–15. Busan, Repub-
lic of Korea. Sponsored by ASM International. www.asminternational.org/content/Events/itsc/. Int’l Conf. on Materials for Renewable Energy & Environment.
May 15, 16. Nanjing, China. www.mree-conf.org . IIE Annual Conf. and Expo. May 18–22. Caribe Hilton, San Juan,
Puerto Rico. www.iienet2.org/annual2. 44th Steelmaking Seminar — Int’l. May 19–22. Tauá Grande
Hotel Termas & Convention Araxá, Estância Parque do Barreiro, s/nº Araxá - Minas Gerais, Brazil. Held by Brazilian Metallurgical, Materials, and Mining Assn. www.abmbrasil.com.br . LPPDE-Europe. June 3–5. Park Plaza Hotel, Amsterdam Air-
port, Amsterdam, Netherlands. Lean Product & Process Development Exchange, Inc. Address e-mail to
[email protected].
bition Centre, Goregaon, Mumbai, India. www.power-genindia. com/index.html.
♦Pipeline Conf. June 4, 5. Houston, Tex. Sponsored by the American Welding Society (800/305) 443-9353, ext. 264; www.aws.org/conferences .
♦ AWS Weldmex Show, FABTEC H Mexi co, METALF ORM
Manufacturing Surabaya 2013. June 12–15. Surabaya, Indonesia.
Mexico. May 7–9. Cintermex, Monterrey, Mexico. Sponsors:
www.pamerindo.com/events/11 . — continued on page 56
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A 360º VIEW OF THE MOST INNOVATIVE TECHNOLOGY AND PROCESSES. FABTECH 2013. METAL FORMING | FABRICATING WELDING | FINISHING FABTECH represents every step of the metal manufacturing process from start to finish. It’s where new ideas, products and technology are highlighted through interactive exhibits, education and networking. Compare solutions from 1,500+ exhibitors, find machine tools to improve quality and productivity, and learn ways to increase profit. REGISTER NOW for the show with a degree of difference.
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18th Beijing-Essen Welding & Cutting Fair. June 18−21. New In-
GAWDA Annual Convention. Sept. 15–18. Orlando, Fla. Gases
ternational Expo Center, Shanghai, China. www.beijing-essen welding.com/en/index.htm.
and Welding Distributors Assn. www.gawda.org. ASM Heat Treating Society Conf. and Expo. Sept. 16–18. Indiana
Third VDI Congress, “Lightweight Design Strategies in Vehicles.” July 3, 4. Wolfsburg, Germany. Sponsored by VDI Wis-
Convention Center, Indianapolis, Ind. www.asminternational.org/ content/Events/heattreat/ .
sensforum, Assn. of German Engineers. www.vdi.de/leichtbau. IIW Int’l Conf. on “Automation in Welding.” Sept. 16, 17. Essen, ♦Codes and Standards Conf. July 16, 17. Orlando, Fla.
To include AWS D1, Structural Welding Code — Steel, ASME Boiler and Pres sure Vessel Code, API pipeline codes, MIL specs and ISO standards. Sponsored by the American Welding Society (800/305) 4439353, ext. 264; www.aws.org/conferences . 59th Annual UA Assn. of Journeymen and Apprentices of the Plumbing and Pipefitting Industry’s Instructor Training Program. Aug. 11–17, Washtenaw Community College, Ann Arbor,
Mich. www.visitannarbor.org/news/detail/ann-arbor-welcomes-the 59th-annual-united-association-instructor-training-p. 12th Int’l Conf. on Application of Contemporary Non-Destructive Testing in Engineering. Sept. 4–6. Grand Hotel Metropol, Por-
toroz, Slovenia. Sponsored by The Slovenian Society for Non-Destructive Testing. www.fs.uni-lj.si/ndt.
Germany. www.iiw2013.com. Event in the IIW Annual Assembly. Schweissen & Schneiden 2013 Int’l Trade Fair — Joining, Cutting, Surfacing. Sept. 16–21. Essen, Germany. Sponsored by DVS (Ger-
man Welding Society). www.schweissenuschneiden.de/en/schweis sen_schneiden/index.html. ♦16th Annual Aluminum Conf.
Sept. 17, 18. Chicago, Ill. Sponsored by the American Welding Society (800/305) 443-9353, ext. 264; www.aws.org/conferences .
Educational Opportunities Brazing School — Fundamentals to Advanced Concepts. May
14–16 (Hartford, Conn.); Oct. 22–24 (Greenville, S.C.); Nov. 19–21 (Simsbury, Conn.). www.kaybrazing.com/s eminars.htm;
[email protected]; (860) 651-5595.
LPPDE-North America. Sept. 9–11. Savannah, Ga. Lean Product
& Process Development Exchange, Inc. Address e-mail to
[email protected].
CWI Preparation Courses: June 3–7, Aug. 19–23, Nov. 11–15. D1.1 Endorsement: June 7, Aug. 23, Nov. 15; D1.5 Endorsement: May 31, Aug. 16; API Endorsement: May 30, Nov. 8. All courses
66th IIW Annual Assembly. Sept. 11–17. Essen, Germany. Or-
and endorsements held at Welder Training & Testing Institute, 1144 N. Graham St., Allentown, Pa. www.wtti.com; (610) 8209551, ext. 204.◆
ganized by DVS (German Welding Society). www.dvs ev.de/IIW2013 /.
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An Association of Welding Manufacturers
Know an individual, company, educator, or educational facility that exemplifies what welding is all about?
Nominate them! The Image of Welding Awards Program recognizes outstanding achievement in the following categories: Individual Individual (you or other individual)
Section Section (A (AWS WS local chapter)
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Entry deadline is July 31, 2013 For mor more e informatio information n and to submit a nomination nomi ation form online, visit www.aws.org/awards/image.html www.aws.org/awards/image.html awards/image.html or call c 800-443-9353.
CERTIFICATION SCHEDULE
Certification Seminars, Code Clinics, and Examinations
Certified Welding Inspector (CWI) LOCATION
SEMINAR D ATES
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Birmingham, AL Hutchinson, KS Spokane, WA Miami, FL Bakersfield, CA Pittsburgh, PA Beaumont, TX Corpus Christi, TX Hartford, CT Orlando, FL Memphis, TN Jacksonville, FL Omaha, NE Cleveland, OH Miami, FL Phoenix, AZ Los Angeles, CA Louisville, KY Waco, TX Milwaukee, WI Corpus Christi, TX Sacramento, CA Kansas City, MO Denver, CO Miami, FL Philadelphia, PA Chicago, IL Baton Rouge, LA Portland, ME Las Vegas, NV Mobile, AL Charlotte, NC Rochester, NY San Antonio, TX Seattle, WA San Diego, CA Minneapolis, MN Salt Lake City, UT Anchorage, AK Miami, FL Idaho Falls, ID St. Louis, MO Houston, TX New Orleans, LA Fargo, ND Pittsburgh, PA Indianapolis, IN
June 2–7 June 2–7 June 2–7 Exam only June 9–14 June 9–14 June 9–14 Exam only June 23–28 June 23–28 June 23–28 July 7–12 July 7–12 July 7–12 Exam only July 14–19 July 14–19 July 14–19 July 14–19 July 14–19 Exam only July 21–26 July 21–26 July 28–Aug. 2 July 28–Aug. 2 July 28–Aug. 2 Aug. 4–9 Aug. 4–9 Aug. 4–9 Aug. 4–9 Aug. 11–16 Aug. 11–16 Exam only Aug. 11–16 Aug. 11–16 Aug. 18–23 Aug. 18–23 Aug. 18–23 Exam only Sept. 15–20 Sept. 15–20 Sept. 15–20 Sept. 15–20 Sept. 22–27 Sept. 22–27 Sept. 22–27 Sept. 29−Oct. 4
June 8 June 8 June 8 June 13 June 15 June 15 June 15 June 29 June 29 June 29 June 29 July 13 July 13 July 13 July 18 July 20 July 20 July 20 July 20 July 20 July 27 July 27 July 27 Aug. 3 Aug. 3 Aug. 3 Aug. 10 Aug. 10 Aug. 10 Aug. 10 Aug. 17 Aug. 17 Aug. 17 Aug. 17 Aug. 17 Aug. 24 Aug. 24 Aug. 24 Sept. 21 Sept. 21 Sept. 21 Sept. 21 Sept. 21 Sept. 28 Sept. 28 Sept. 28 Oct. 5
9–Year Recertification Seminar for CWI/SCWI (No exams given.) For current CWIs and SCWIs needing to meet education requirements without taking the exam. The exam can be taken at any site listed under Certified Welding Inspector. LOCATION
SEMINAR D ATES
Pittsburgh, PA San Diego, CA
June 2–7 July 7–12
Miami, FL Orlando, FL Denver, CO Dallas, TX
July 21–26 Aug. 18–23 Sept. 15–20 Oct. 6–11
Certified Welding Supervisor (CWS) LOCATION
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Minneapolis, MN Miami, FL Norfolk, VA
July 15–19 Sept. 23–27 Oct. 14–18
July 20 Sept. 28 Oct. 19
CWS exams are also given at all CWI exam sites. Certified Radiographic Interpreter (CRI) LOCATION
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Dallas, TX Chicago, IL Pittsburgh, PA
Aug. 19–23 Sept. 23–27 Oct. 14–18
Aug. 24 Sept. 28 Oct. 19
The CRI certification can be a stand-alone credential or can exempt you from your next 9-Year Recertification. Certified Welding Sales Representative (CWSR) CWSR exams will be given at CWI exam sites. Certified Welding Educator (CWE) Seminar and exam are given at all sites listed under Certified Welding Inspector. Seminar attendees will not attend the Code Clinic portion of the seminar (usually the first two days). Certified Robotic Arc Welding (CRAW) The course dates are followed by the location and phone number
June 17–21, Dec. 9–13 at ABB, Inc., Auburn Hills, MI; (248) 391–8421
May 20–24, Aug. 19–23, Dec. 2–6 at Genesis-Systems Group, Davenport, IA; (563) 445-5688
Oct. 14 at Lincoln Electric Co., Cleveland, OH; (216) 383-8542
July 15–19, Oct. 21–25 at OTC Daihen, Inc., Tipp City, OH; (937) 667-0800
Training: May 20–22, July 22–24, Sept. 23–25, Nov. 18–20 Exams: May 23–24, July 25–26, Sept. 26–27, Nov. 21–22 at Wolf Robotics, Fort Collins, CO; (970) 225-7736
On request at MATC, Milwaukee, WI; (414) 297-6996
Certified Welding Engineer; Senior Certified Welding Inspector Exams can be taken at any site listed under Certified Welding Inspector. No preparatory seminar is offered. International CWI Courses and Exams Schedules Please visit www.aws.org/certification/inter_contact.html.
IMPORTANT: This schedule is subject to change without notice. Applications are to be received at least six weeks prior to the
seminar/exam or exam. Applications received after that time will b e assessed a $250 Fast Track fee. Please verify application deadline dates by visiting our website www.aws.org/certification/docs/schedules.html . Verify your event dates with the Certification Dept. to confirm your course status before making travel plans. For information on AWS seminars and certification programs, or to register online, visit www.aws.org/certification or call (800/305) 443-9353, ext. 273, for Certification; or ext. 455 for Seminars. Apply early to avoid paying the $250 Fast Track fee. 58
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METALLURGY for the Non-Metallurg Non-Metallurgist: ist: Fundamentals Metallurgy is the science that deals with the iinternal nternal structure of metals, the relationship between metals, and the properties of metals. In welding, a basic understanding of metallurgy provides insight into the positive and negative changes that occur in metals when joined by welding. From the properties of an atom to the behavi behaviors ors of metals during the welding process, you are introduced to the properties of metals and will gain an understanding of why metals behave the way they do. Concepts covered include the anatomy of ato atoms, ms, the periodic table, chemical bonding, including ionic bonding, covalent bonding, and metallic bonding, as well as the properties of metals. This seminar contains interactive exercises to reinforce key points and includes summarie and quizzes to help prepare you for the completion exam. summaries The seminar is approximately five hours long and concludes with a proficiency test.
Sample seminar at awo.aws.org/seminars/metallurgy
CONFERENCES
Pipeline Conference June 4, 5 Houston, Tex. The transportation of oil and natural gas through cross-country pipelines has never been as vigorous as it is now, and greater growth lies ahead with welding in the thick of it. For many decades, the stick electrode has been a driving force behind the construction of these lines, and it is still very much in the driver’s seat. But in order to cut costs, owners have started to use X80, a lighter-weight, higher-strength linepipe steel. The same cellulosic electrodes used on the more conventional steels are inadequate for X80 steel. This has opened the door for low-hydrogen electrodes and mechanized welding. The keynote address to this important AWS conference will be delivered by Brian Laing, president of CRC-Evans Pipeline International. A seasoned veteran of the pipeline industry, Laing once worked for NOVA (now TransCanada), as a welding engineer. Following is a breakdown of the other speakers and their topics. Robin Gordon, senior vice president of Microalloying International Inc., will present the current status of Grade X80 pipeline technology and highlight the technical challenges that must be addressed before considering its use. Paul Tews’s presentation o n “Specimen Quality for Fatigue Test Girth Welds” should be of interest to every pipeline owner. Tews, who is operating out of the UK at present, is the principal
welding engineer for Subsea 7. Bill Bruce, U.S. director of Welding and Materials Technology for Det Norske Veritas, will discuss new revisions for pipeline repair in API 1104. Two popular processes, hybrid laser arc welding and friction stir welding, are waiting their turns for acceptance in some applications. Matt Boring, senior welding engineer, Knieper & Associates, will speak on the situation from the standpoint of ASME Section IX. Ian Harris, technical leader, Arc Welding, EWI, will give a detailed presentation on the hybrid laser arc welding process and how it is suited for pipeline construction work. Another speaker from EWI, Connie Reichert, will talk about automated corrosion repair of pipelines. Reichert is principal engineer, Design, Controls & Automation. Michael Lang, senior construction engineer, Bechtel Corp., has lean welding as his topic. Lang is chairman of the AWS D10 Committee on Piping and Tubing. Russel Fuchs, senior technical manager, Bohler Welding Group USA, will offer a comparison of one cellulosic electrode and two low-hydrogen electrodes. “SMAW: The Evo lution of Stick Welding, from a Welder’s Perspective” is the title of Lori Kuiper’s talk. She is offshore & pipeline segment manager, Euroweld, Ltd. Derick Railling, product manager, Global Onshore Pipeline, ITW Welding, is basing his presentation on understanding the sources and remedies of hydrogen-induced cracking in pipeline welds. Scott Funderburk from CRC-Evans Pipeline International will talk about overcoming such o perational challenges as leak detection and automatic shut-off valves. Chris Penniston, welding and materials engineer, RMS Systems in Canada, has chosen the topic of innovations in mechanized welding. Olivier Jouffron, technical manager, Serimax North America, will discuss the steps that can be taken in welding corrosionresistant alloy pipe. Win Wijnholds, president of Magnatech International BV in The Netherlands, will discuss dual-process methodology. Ryan Lewis, a consumables product manager at The Lincoln Electric Co., will discuss some of the welding activities used in oil shale environments.
Codes and Standards Conference July 16, 17 Orlando, Fla. This conference will feature information about the AWS D1 Structural Welding Code — Steel , ASME Boiler and Pressure Vessel Code, and API pipeline codes, plus M IL and ISO standards, potentially the most valuable documents available to manufacturers and fabricators of welded products. Information will be provided about the planning and execution of various welding processes, as well as useful data for designers, inspectors, and QC specialists. ◆
For more information, please contact the AWS Conferences and Seminars Business Unit at (800) 443-9353, ext. 223, or e-mail
[email protected] . You can also visit the Conference Department Web site at www.aws.org/conferences for upcoming conferences and registration information. For info go to www.aws.org/ad-index
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WELDING WORKBOOK
Datasheet 340
Selecting Shielding Gases for Gas Metal Arc Welding The primary function of the shielding gas in gas metal arc welding (GMAW) is to exclude the atmosphere from contact with the molten weld metal. This is necessary because most metals, when heated to their melting point in air, exhibit a strong tendency to form oxides, and, to a lesser extent, nitrides. Oxygen also reacts with carbon in molten steel to form carbon monoxide and carbon dioxide. These reaction products may result in weld discontinuities such as slag inclusions, porosity, and weld metal embrittlement. Reaction products are easily formed in the at-
mosphere unless precautions are taken to exclude nitrogen and oxygen. Besides providing a protective environment, the shielding gas and flow rate have a pronounced effect on arc characteristics, mode of metal transfer, penetration and weld bead profile, speed of welding, undercutting tendency, cleaning action, and weld metal mechanical properties. The principal gases used in the spray arc mode are shown in Table 1 and those for short-circuit GMAW are shown in Table 2. ◆
Table 1 — GMAW Shielding Gases for Spray Transfer Metal
Shielding Gas
Characteristics
Aluminum
100% argon 35% argon/65% helium 25% argon/75% helium
Best metal transfer and arc stability; least spatter; good cleaning action. Higher heat input than 100% argon; improved fusion characteristics on thicker material; minimizes porosity. Highest heat input; minimizes porosity; least cleaning action.
Magnesium
100% argon Argon + 20–70% helium
Excellent cleaning action; stable arc. Improved wetting; less chance of porosity.
Carbon steel
1–5% oxygen, balance argon
Improves arc stability; produces a more fluid and controllable weld pool; good fusion and bead contour; minimizes undercutting; permits higher speeds than pure argon. High-speed mechanized welding; low-cost manual welding.
5–15% carbon dioxide (CO 2); balance argon Low-alloy steel
98% argon/2% oxygen
Minimizes undercutting; provides good toughness.
Stainless steel
99% argon/1% oxygen
Improves arc stability; produces a more fluid and controllable weld pool; good fusion and bead contour; minimizes undercutting on heavier stainless steels. Provides better arc stability, coalescence, and welding speed than 1% oxygen mixture for thinner stainless steel materials.
98% argon/2% oxygen Nickel, copper, and their alloys
100% argon Argon/helium
Provides good wetting; decreases fluidity of weld metal. Higher heat inputs of 50 and 75% helium mixtures offset high heat dissipation of heavier gauges.
Titanium
100% argon
Good arc stability; minimum weld contamination; inert gas backing is required to prevent air contamination on back of weld area.
Table 2 — GMAW Shielding Gases for Short-Circuiting Transfer Metal
Shielding Gas
Characteristics
Carbon Steel
75% argon/25% CO2
High welding speeds with minimum melt-through; minimum spatter; clean weld appearance; good pool control in vertical and overhead positions. Deep penetration; faster welding speeds; high spatter levels.
100% CO2 Stainless steel
90% helium/7.5% argon/2.5% CO2
No effect on corrosion resistance; small heat-affected zone; minimizes undercut.
Low-alloy steel
60–70% helium/25–35% argon/ 4.5% CO2 75% argon/25% CO2
Minimum reactivity; excellent toughness; excellent arc stability, wetting characteristics, and bead contour; little spatter. Fair toughness; excellent arc stability, wetting characteristics, and bead contour; little spatter.
Aluminum, copper magnesium, nickel, and their alloys
Argon and argon/helium
Argon satisfactory on sheet metal; argon-helium preferred for thicker base material.
Excerpted from the Welding Handbook, Vol. 2, ninth edition. 62
MAY 2013
AWS Conferences & Exhibitions:
Pipelines Conference June 4th – 5th / Houston, TX Join us in Houston for the debut of the AWS Pipeline Welding Conference! Our featured speakers will cover a multitude of topics including the welding of high strength X80 pipe steels, orbital processes used in pipeline construction throughout the world, the new FRIEX system from Belgium and many other exciting topics. Highlights
Learn about the progress of new and innovative developments in pipeline welding. Network with industry peers to �nd the best solutions for business growth.
AWS Conference attendees are awarded 1 PDH (Professional Development Hour) for each hour of conference attendance. These PDH's can be applied toward AWS recerti�cations and renewals.
For the latest conference information and registration visit our web site at www.aws.org/conferences or call 800-443-9353, ext. 224.
SOCIETYNEWS BY HOWARD WOODWARD
[email protected]
D1 Committee Convenes in Doral
The D1 Committee on Structural Welding met at AWS World Headquarters the week of Feb. 25.
Shown are (from left) AWS Vice President David McQuaid, past D1 Chair Donald Rager, AWS Executive Director Ray Shook, D1 Chair Duane Miller, and Allen Sindel, D1 vice chair.
On March 1, a rare copy of the first-edition of the forerunner to the D1.1, Structural Welding Code — Steel , all 20 pages of it, was presented to Ray Shook , AWS executive director, for permanent loan and preservation in the D. Fred Bovie Library and Museum at AWS World Headquarters in Doral, Fla. The presenters included Duane K. Miller, D1 chair; David McQuaid, D1 chair (1996–2002)
and AWS vice president; Donald D. Rager, D1 chair (2002–2008), and Allen W. Sindel, D1 vice chair. The Code for Fusion Welding and Gas Cutting in Building Construction includes general application, definitions, materials, permissible unit stresses, design, workmanship, erection, gas cutting, and an appendix. The document was donated to the Society by Paul E. Masters, Cape Coral, Fla. WELDING JOURNAL
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AWS President Nancy Cole (left) presents Anne Rorke a special AWS award for her many years of service at WTIA (Welding Tech nology Institute of Australia) where she cur rently is editor of the Australasian Welding Journal . The presentation was made at the March WTIA conference in Perth, Australia.
David McQuaid (left), an AWS vice president and a past chair of the D1 Committee, re ceives recognition for his 30 years of service from D1 Committee Chair Duane Miller.
Damian J. Koteck i (r ight) receives his 30 year service award pin from Harry Wehr, chairman, A5 Committee on Filler Metals and Allied Materials, on March 5, during the annual me etings in Orlando, Fla. Kot ecki was cited for his many contributions to the A5 Committees.
Tech Topics New Standards Projects Development work has begun on the following revised standards. Affected indi viduals are invited to cont ribute to their development. Contact Staff Secretary A. Diaz, adiaz @aws.org ; ext. 304. Participation on AWS Technical Committees is open to all persons. B2.1-1/8-010:20XX, Standard Welding Procedure Speci fication (SWPS) for Gas Tungsten Arc Welding of Carbon Steel to Austenitic Stainless St eel (M-1, P-1, or S -1 to M-8, P-8, or S-8), 18 through 10 Gauge, in the As-Welded Condition, with or without Backing. Stakeholders: Manufacturers, welders, CWIs, engineers. B2.1-1/8-231:20XX, Standard Welding Procedure Speci fication (SWPS) for Gas Tungsten Arc Welding with Consumable In sert Root followed by Shielded Metal Arc Welding of Carbon Steel (M-1/P-1/S-1, Groups 1 or 2) to Austenitic Stainless Steel 1 1 8 through 1 ⁄ 2 Inch (M-8/P-8/S-8, Group 1), ⁄ Thick, IN309, ER309, and E309-15,-16, or -17, or IN309, ER309(L), and ER309(L)15, -16, or -17, As-Welded Condition, Prima rily Pipe Applications. Stakeholders: Manufacturers, welders, CWIs, and welding engineers. D9.1M/D9.1:20XX , Sheet Metal Welding Code. Stakeholders: Those involved in the production and qualification of nonstructural sheet metal applications such as heating, ventilating, and air conditioning systems, food processing equipment, architectural sheet metal, and in the acceptance of welding and braze welding of nonstructural sheet metal components.
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ANSI Approved Revised Standards A5.16/A5 .16M:201 3 (ISO 2403 4:20 10 MOD) , Specification for Titanium and Ti-
tanium-Alloy Welding Electrodes and Rods. Approved 2/19/13. F1.2:2013, Laboratory Method for Measuring Fume Generation Rates and Total Fume Emission of Welding and Allied Processes. Approved 2/25/13. ANSI Approved Reaffirmed Standards B2.1-1-003:2002 (R2013) , Standard Welding Procedure Specification (SWPS) for Gas Metal Arc Welding (Short Circuiting Transfer Mode) of Galvanized Steel (M-1), 18 through 10 Gauge, in the As-Welded Con dition, with or without Backing. 3/7/13. B2.1-1-004:2002 (R2013) , Standard Welding Procedure Specification (SWPS) for Gas Metal Arc Welding (Short Circuiting Transfer Mode) of Carbon Steel (M-1, Group 1), 18 through 10 Gauge, in the As-Welded Condition, with or without Backing. 3/7/13. B2.1-8-005:2002 (R2013) , Standard Welding Procedure Specification (SWPS) for Gas Metal Arc Welding (Short Circuiting Transfer Mode) of Austenitic Stainless Steel (M-8, P-8, or S-8), 18 through 10 Gauge, in the As-Welded Condition, with or without Backing. 3/7/13. B2.1-1/8-006:2002 (R2013) , Standard Welding Procedure Specification (SWPS) for Gas Metal Arc Welding (Short Circuiting Transfer Mode) of Carbon Steel to Austenitic Stainless Steel (M-1 to M-8, P-8, or S-8), 18 through 10 Gauge, in the As-Welded Condition, with or without Backing. 3/14/13. B2.1-1-007:2002 (R2013) , Standard Welding Procedure Specification (SWPS) for
Gas Tungsten Arc Welding of Galvanized Steel (M-1), 18 through 10 Gauge, in the AsWelded Condition, with or without Backing. 3/7/13. B2.1-1-008:2002 (R2013) , Stand ard Welding Procedure Specification (SWPS) for Gas Tungsten Arc Welding of Carbon Steel (M-1, P-1, or S-1), 18 through 10 Gauge, in the As-Welded Condition, with or without Backing. 3/11/13. B2.1-8-009:2002 (R2013) , Stand ard Welding Procedure Specification (SWPS) for Gas Tungsten Arc Welding of Austenitic Stainless Steel (M-8, P-8, or S-8), 18 through 10 Gauge, in the As-Welded Condition, with or without Backing. 3/11/13. B2.1-1-011:2002 (R2013) , Stand ard Welding Procedure Specification (SWPS) for Shielded Metal Arc Welding of Galvanized Steel (M-1), 10 through 18 Gauge, in the AsWelded Condition, with or without Backing. 3/11/13. B2.1-1-012:2002 (R2013) , Stand ard Welding Procedure Specification (SWPS) for Shielded Metal Arc Welding of Carbon Steel (M-1, P-1, or S-1), 18 through 10 Gauge, in the As-Welded Condition, with or without Backing. 3/11/13. B2.1-8-013:2002 (R2013) , Stand ard Welding Procedure Specification (SWPS) for Shielded Metal Arc Welding of Austenitic Stainless Steel (M-8, P-8, S-8, Group 1), 10 through 18 Gauge, in the As-Welded Condition, with or without Backing. 3/11/13. B2.1-1/8-014:2002 (R2013) , Stand ard Welding Procedure Specification (SWPS) for Shielded Metal Arc Welding of Carbon Steel to Austenitic Stainless Steel (M-1 to M-8/P8/S-8, Group 1), 10 through 18 Gauge, in the
As-Welded Condition, with or without Backing. 3/11/13. B2.1-1/8-227:2002-AMD1 (R2013) , Standard Welding Procedure Specification (SWPS) for Gas Tungsten Arc Welding of Carbon Steel (M-1/P-1, Groups 1 or 2) to Austenitic Stainless Steel ( M-8/P-8, Group 1 1 16 through 1 ⁄ 2 Inch Thick, ER309(L), 1), ⁄ As-Welded Cond ition, Primarily Pipe Ap plications. 3/11/13. B2.1-1/8-228:2002 (R2013) , Standard Welding Procedure Specification (SWPS) for Shielded Metal Arc Welding of Carbon Steel (M-1/P-1/S-1, Groups 1 or 2) to Austen itic Stainless Steel ( M-8/P-8/S -8, 1 1 8 through 1 ⁄ 2 Inch Thick, Group 1), ⁄ E309(L) -15, -16, or -17, As-Welded Con dition, Primarily Pipe Applications. 3/11/13. B2.1-1/8-229:2002-AMD1 (R2013) , Standard Welding Procedure Specification (SWPS) for Gas Tungsten Arc Welding fol lowed by Shielded Metal Arc Welding of Carbon Steel (M-1/P-1, Groups 1 or 2) to Austenitic Stainless Steel (M-8/P-8, Group 1 1 8 through 1 ⁄ 2 Inch Thick, ER309(L) and 1), ⁄ E309(L) -15, -16, or -17, As-Welded Con dition, Primarily Pipe Applications. 3/11/13. B2.1-1/8-230:2002-AMD1 (R2013) , Standard Welding Procedure Specification (SWPS) for Gas Tungsten Arc Welding with Consumable Insert Root of Carbon Steel (M-1/P-1, Groups 1 or 2) to Austenitic 1 16 Stainless Steel (M-8/P-8, Group 1), ⁄ 1 through 1 ⁄ 2 Inch Thick, IN309 and
ER309(L), As-Welded Conditi on, Prima rily Pipe Applications. 3/11/13. Standards for Public Review B5.17:20XX, Specification for the Qualification of Welding Fabricators . Revised.
$25. 5/20/13. C1.5:20XX, Specification for the Quali-
fication of Resistance Welding Technicians.
are available from your national standards body, which in the United States is ANSI, 25 W. 43rd St., 4th Fl., New York, NY, 10036; (212) 642-4900. Send comments regarding ISO documents to your national standards body. In the United States, if you wish to participate in the development of International Standards for welding, contact A. Davis,
[email protected].
Revised. $25. 5/20/13. D15.1/D15.1M:2012-AMD1 , Railroad
Welding Specification for Cars and Loco motives. Amendment standard. $129.
5/13/13. 2nd BSR-8. AWS was approved as an accredited standards-preparing organization by the American National Standards Institute (ANSI) in 1979. AWS rules, as approved by ANSI, require that all standards be open to public review for comment during the approval process. The above standards are submitted for public review with the closing dates shown. A draft copy may be obtained from R. O’Neill,
[email protected] . ISO Standard for Public Review ISO/DIS 5826.2 , Resistance weldi ng equipment — Transformers — General specifications applicable to all transformers ISO/DIS 17533 , Welding for aerospace applicat ions — Weldin g da ta in de sign documents
Technical Committee Meetings
All AWS technical committee meetings are open to the public. Persons wishing to attend a meeting should contact the committee secretary listed. May 1, D8 Committee on Automotive Welding. Livonia, Mich. E. Abrams, ext. 307. May 7–9, D17 Committee on Welding in the Aircraft and Aerospace Industries. Los Angeles, Calif. A. Diaz, ext. 304. May 16 , Safety and Health Committee. Cleveland, Ohio. S. Hedrick, ext. 305. May 28–30 , D14 Committee on Machinery and Equipment. Dallas, Tex. E. Abrams, ext. 307. June 5, C2 Committee on Thermal Spraying. Ogden, Utah. E. Abrams, ext. 307. June 18, G2D Subcommittee on Reactive Alloys. Seattle, Wash. A. Diaz, ext. 304.
Review copies of the above documents
Share Your Technical Expertise Volunteers are sought to contribute to the following technical committees Visit www.aws.org/technical/jointechcomm.html AWS Safety and Hea lth Com mit tee
seeks educators, users, general interest, and consultants to help develop standards on welding safety. S. Hedrick,
and Cutting, and D8 Committee on Automotive Welding seek educators, general
D16 Committee on Robotic and Automatic Welding seeks members in the gen-
interest, and end users to update its documents. E. Abrams,
[email protected] .
eral interest and educational fields to help revise its documents. B. McGrath, bmc
[email protected] .
[email protected]. B4 Committee on Mechanical Testing of Welds seeks professionals in the area
of standard methods for tension, shear, bend, fracture toughness, hardness, weldability, and other mechanical testing of welds. B. McGrath,
[email protected]. B2B Subcommittee on Welding Qualifications , seeks members to update B2.1,
A5L Subcom mit tee on Magnes ium Alloy Filler Metals seeks professionals to
revised its filler metal document. R. Gupta,
[email protected] . D10P Subcommittee for Local Heat Treating of Pipe seeks heat treating pro-
fessionals to help update its documents. B. McGrath,
[email protected] .
Specification for Welding Procedure and Perfo rmance Qua lifica tio n . A. Diaz,
[email protected] .
D14 Committee on Machinery and Equipment and D14H Subcommittee on Surfacing and Reconditioning of Industrial Mill Rolls seeks professionals in de-
D17J Subcommittee seeks profession-
sign, production, engineering, testing, and safe operation of machinery to prepare recommended practices for surfacing and reconditioning of industrial mill rolls. The next meeting of the D14 Committee is May 28 in Dallas, Tex. To attend, contact E. Abrams,
[email protected] .
als to update specification for friction stir welding of aluminum alloys for aerospace applications. A. Diaz,
[email protected] . C2 Committee on Thermal Spraying, C4 Committee on Oxyfuel Gas Welding
G2D Subcommittee on Reactive Alloys
seeks volunteers to update guides for the fusion welding of titanium and titanium alloys, and fusion welding of zirconium and zirconium alloys. A. Diaz,
[email protected] . J1 C ommitt ee o n Res ist ance Welding Equipment seeks educators, general in-
terest, and users to develop standards on controls, installation and maintenance, calibration, and resistance welding fact sheets. E. Abrams,
[email protected] . A5K Sub com mit tee on Tit ani um and Zirconium Filler Metals . Seeks profes-
sionals in the field to update specifications for welding electrodes and rods of titanium, zirconium, and their alloys. A. Diaz,
[email protected] .
WELDING JOURNAL
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Three Student Chapter Advisors Announce Awards Roy Ledford Jr. , advisor, Lawson State Community College Student Chapter, Bessemer, Ala., has selected Benjamin Vining and Randall Standridge to receive the Student Chapter Member Award. Vining , who is Chapter treasurer, will represent the college at the 2013 Louisiana State SkillsUSA competition. He has passed several welding procedure qualifications while maintaining a 4.0 GPA. He has been named the Outstanding Student in the Welding Technology Program, and is named on this year’s President’s List. Standridge, 2012– 13 Student Chapter chair, represented the Chapter at the “Meet the President” event. He serves as a volunteer tutor for several subjects, has directed student activities, and coordinated field trips and other events.
William Burns, advisor, Savannah Technical College “Southern Welders” Student Chapter, Savannah, Ga., has selected Dustin Bolgrign to receive the 2012–13 Student Chapter Member Award. Bolgrign served as the 2012–13 Chapter chair and earned a 3.95 GPA in his final semester in the welding program. He served nine years in the U.S. Army, completing four combat tours, and continues to serve in the Army Reserves. Huck Hughes, advisor, Columbiana County Career and Technical Center Student Chapter, Lisbon, Ohio, has selected Paige Lawrence , Chapter chair, to receive the Student Chapter Member Award. The AWS Board of Directors established the Student Chapter Member Award
to recognize AWS Student Members whose Student Chapter activities have produced outstanding school, community, or industry achievements. This award also provides an opportunity for Student Chapter advisors, Section officers, and District directors to recognize outstanding students affiliated with AWS Student Chapters, as well as to enhance the image of welding within their communities. To qualify for this certificate award, the individual must be an AWS Student Member affiliated with an AWS Student Chapter. The criteria and nomination form can be downloaded from www.aws.org/ sections/awards/student_chapter.pdf , or request a copy from the Membership Dept. (800) 443-9353, ext. 260.
Name Your Candidates for These AWS Awards The deadline for nominating candidates for the following awards is December 31 prior to the year o f the awards presentations. Contact Wendy Sue Reeve,
[email protected]; (800/305) 443-9353, ext. 293. William Irrgang Memorial Award
is given to the individual who has done the most over the past five years to enhance the Society’s goal of advancing the science and technology of welding. It includes a $2500 honorarium and a certificate. This award
Honorary Membership Award
This award acknowledges eminence in the welding profession, or one who is credited with exceptional accomplishments in the development of the welding art. Honorary Members have full rights of membership. Nat. Meritorious Certificate Award
This award recognizes the recipient’s counsel, loyalty, and dedication to AWS affairs, assistance in promoting cordial rela-
tions with industry and other organizations, and for contributions of time and effort on behalf of the Society. George E. Willis Award
This award is given to an individual who promoted the ad vancemen t of welding i nte rna tio nal ly by fostering coopera tive participation in technology transfer, standards rationalization, and promotion of industrial goodwill. It includes a $2500 honorarium. International Meritorious Certificate Award
This honor recognizes recipients’ significant contributions to the welding industry for service to the international welding community in the broadest terms. The award consists of a certificate and a one-year AWS membership.
New Award Category Created: AWS Distinguished Welder The AWS Distinguished Welder Award has been created to recognize professionals with a minimum of 15 years’ experience as a welder and/or supervisor whose welding skills and experience warrant this special recognitio n. August 1 is the deadline for submitting your completed nomination form. The nomination packet should include information addressing the Definition and candidate’s Application Criteria as detailed in the AWS Distinguished Welder Award Nomination Form. The focus of the nomination packet should include specifics of the individual’s skills. For details on the full description, selection criteria, and nomination form, visit the AWS Web site
and select the awards category or e-mail Wendy Sue Reeve, senior manager, awards programs,
[email protected] . A maximum of ten individuals per year may be selected as Distinguished Welders as determined by the Selection Committee. Nominations shall remain valid for three years. If the maximum number of Distinguished Welders allowed under the rules is reached, the remaining candidates will be deferred for consideration the next year, consistent with their time eligibility. If less than the maximum number are awarded, the remaining candidates will be deferred for consideration the following year, consistent with their time eligibility.
Candidates Sought for Annual Masubuchi Award November 1, 2013, is the deadline for submitting nominations for the 2014 Prof. Koichi Masubuchi Award. This award includes a $5000 honorarium. It is presented each year to one person, 40 years old or younger, who has made significant contributions to the advancement of materials joining through research and de velopment. Nominations should include a description of the can-
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didate’s experience, list of publications, honors, and awards, and at least three letters of recommendation from fellow researchers. The award is sponsored by the Massachusetts Institute of Technology Dept. of Ocean Engineering. E-mail your nomination package to Todd A. Palmer, assistant professor, The Pennsylvania State University,
[email protected].
New AWS Supporters
New Sustaining Members Babcock & Wilcox Nuclear Energy, Inc. 11525 N. Community House Rd. Charlotte, NC 28277 Representative: Daniel E. Applegate www.babcock.com
Bluescope Building North America, Inc. 7440 Doe Ave. Visalia, CA 93291 Representative: Thomas Andersen www.bluescopesteel.com
Greenville Technical College 738 S. Pleasantburg Dr. Greenville, SC 29607 Representative: Jerry D. Norris www.gvltec.edu Greenville Tech’s welding program delivers a strong combination of hands-on experience with classroom instruction. Its graduates are fully trained to enter the work force and are prepared to adva nce into supervisory positions.
Affiliate Companies Big B Welding Service 64016 Arcola Railroad Ln. Roseland, LA 70456
Educational Institutions ARPEC (Air Conditioning, Refrigeration & Pipefitting Education Center)
13201 NW 45th Ave. Opa-locka, FL 33054
Cab Construction Co.
1532 1st Ave. Mankato, MN 56001 ESGA Ingenieria en Estructura S.A. de C.V., Prolongacior de Recursos
Hidraulicos #6 La Loma Tlalnepantla 54060, Mexico
Birmingham Ironworkers Training Program Trust
2828 4th Ave. S. Birmingham, AL 35233 Carrollton Area Career Center
305 E. 10th St. Carrollton, MO 64633
FM Stainless, LLC
1524 Ray Mountain Rd. Ellijay, GA 30536
Cerro Coso Community College
3000 College Height Blvd. Ridgecrest, CA 93555
JRV Construction Enterprises, Inc.
891 Aura Rd. Glassboro, NJ 08028
Mastbaum AVTS High School
3116 Frankford Ave. Philadelphia, PA 19134
Leading Edge Mfg.
303 Chemin Metairie Rd. Youngsville, LA 70592
Midlands Technical College
1260 Lexington Dr. West Columbia, SC 29170
McMenimen Design and Fabrication
3100 Cedar Bay Dr. Melbourne, FL 32934
Muskegon Community College
221 S. Quarterline Rd. Muskegon, MI 49442
Webb Diving Services
Linear Controls, Inc. 1 2 Commission Blvd. 107 ⁄ Lafayette, LA 70508 Representative: Mckenna Bergeron www.linearcontrols.net
Multi-Contact USA 100 Market St. Windsor, CA 95492 Representative: Jocelyn Owen www.multi-contact-usa.com
6409 Rutledge Pike Knoxville, TN 37924
Nueces Canyon High School
200 Taylor St. Barksdale, TX 78828
Westar d.b.a. Quik-Shor
13217 Laureldale Ave. Downey, CA 90242
Pitt Community College
2064 Warren Dr. Winterville, NC 28590
Supporting Companies Al Yousuf Enterprises
Tennessee Technology Center at Crump
215, Aisha Bldg., Office No. 1 SVP Rd., Dongri, Mumbai Maharashtra 400009, India
3070 Hwy. 64, PO Box 89 Crump, TN 38327 White Deer ISD - Agriculture Dept.
Hickey Metal Fabrication
Pulverman 1170 Lower Demunds Rd. Dallas, PA 18612 Representative: Scott Stephenson www.pulverman.net
The Japan Welding Engineering Soc. Qualification & Certification Dept. 4-20 Kanda Sakuma-Cho,Chiyoda-Ku Tokyo 101-0025, Japan Representative: Masaharu Sato www.jwes.or.jp
873 Georgetown Rd. Salem, OH 44460
604 Doucette White Deer, TX 79097 Welding Distributors
H & R Welding, LLC
ACIT (USA), Inc.
307 Drum Point Rd. Brick, NJ 08723
6333 Hazelwood Ln. SE Bellevue, WA 98006
Sureway Tool & Engineering Co.
Consolidated Steel Services, Inc.
2959 Hart Dr. Franklin Park, IL 60131
632 Glendale Valley Blvd. Fallentimber, PA 16639
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Member-Get-A-Member Campaign Listed are the members participating in D. Saunders, Lakeshore — 3 the 2012−2013 campaign. Standings as of J. Turcott, Rochester — 3 March 18, 2013. See page 83 of this Weld- A. Winkle, Kansas City — 3 ing Journal for campaign rules and prize list R. Wright, San Antonio — 3 or visit www.aws.org/mgm. For information, R. Zabel, SE Nebraska — 3 call the Membership Department President’s Honor Roll (800/305) 443-9353, ext. 480. Winner’s Circle Sponsored 2 Individual Members Sponsored 20 or more Individual Mem- G. Cornell, St. Louis bers per year since June 1, 1999. The super- M. Depuy, Portland script denotes the number of times the mem- D. Hayes Jr., Louisville ber achieved this status if more than once. J. Helfrich, Tri-River E. Ezell, Mobile10 P. Host, Chicago 7 J. Compton, San Fernando Valley H. Hughes, Mahoning Valley J. Merzthal, Peru 2 J. Kline, Northern New York G. Taylor, Pascagoula 2 L. Kvidahl, Pascagoula 2 L. Taylor, Pascagoula W. Larry, Southern Colorado B. Chin, Auburn G. Lawrence, N. Central Florida S. Esders, Detroit J. Mansfield, Philadelphia M. Haggard, Inland Empire E. Norman, Ozark M. Karagoulis, Detroit A. Sam, Trinidad S. McGill, NE Tennessee D. Saunders, Lakeshore B. Mikeska, Houston C. Shepherd, Houston W. Shreve, Fox Valley G. Solomon, Central Pennsylvania T. Weaver, Johnstown/Altoona A. Sumal, British Columbia G. Woomer, Johnstown/Altoona C. Villarreal, Houston R. Wray, Nebraska J. Vincent, Kansas City A. Vogt, New Jersey President’s Guild J. Vorstenbosch, International Sponsored 20+ new Individual Members M. Wheeler, Cleveland M. Pelegrino, Chicago — 30 L. William, Western Carolina E. Ezell, Mobile — 22 W. Wilson, New Orleans J. Winston, St. Louis President’s Roundtable Sponsored 9–19 new Individual Members Student Member Sponsors Sponsored 4+ new Student Members R. Fulmer, Twin Tiers — 10 W. Blamire, Atlanta — 9 H. Hughes, Mahoning Valley — 106 A. Tous, Costa Rica — 9 A. Theriot, New Orleans — 47 P. Strother, New Orleans — 9 B. Scherer, Cincinnati — 39 D. Saunders, Lakeshore — 36 President’s Club W. England, Western Michigan — 33 Sponsored 3–8 new Individual Members R. Bulthouse, Western Michigan — 31 D. Galigher, Detroit — 7 R. Hammond, Greater Huntsville — 28 W. Komlos, Utah — 7 T. Geisler, Pittsburgh — 24 J. Smith, San Antonio — 6 S. Siviski, Maine — 24 C. Becker, Northwest — 5 R. Zabel, SE Nebraska — 24 L. Webb, Lexington — 4 B. Cheatham, Columbia — 23 D. Wright, Kansas City — 4 C. Kochersperger, Philadelphia — 23 T. Baber, San Fernando Valley — 3 M. Arand, Louisville — 22 J. Bain, Mobile — 3 D. Bastian, NW Pennsylvania — 21 A. Bernard, Sabine — 3 G. Gammill, NE Mississippi — 21 J. Blubaugh, Detroit — 3 F. Oravets, Pittsburgh — 20 P. Brown, New Orleans — 3 J. Theberge, Boston — 20 D. Buster, Eastern Iowa — 3 J. Johnson, Madison-Beloit — 19 C. Daon, Israel Section — 3 V. Facchiano, Lehigh Valley — 18 G. Gammill, NE Mississippi — 3 J. Falgout, Baton Rouge — 18 B. Hackbarth, Milwaukee — 3 R. Munns, Utah — 18 S. Jaycox, Long Island — 3 S. Lindsey, San Diego — 17 D. Jessop, Mahoning Valley — 3 R. Richwine, Indiana — 17
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AWS Member Counts April 1, 2013 Sustaining ......................................573 Supporting.....................................335 Educational ...................................619 Affiliate..........................................503 Welding Distributor........................53 Total Corporate ..........................2,083 Individual .................................58,817 Student + Transitional .................9,471 Total Members .........................68,288
J. Russell, Fox Valley — 17 M. Anderson, Indiana — 16 E. Norman, Ozark — 16 M. Anderson, Indiana — 16 C. Donnell, NW Ohio — 14 R. Hutchinson, Long Bch./Or. Cty. — 14 D. Pickering, Central Arkansas — 13 C. Daily, Puget Sound — 12 J. Daugherty, Louisville — 12 C. Morris, Sacramento — 12 S. Robeson, Cumberland Valley — 12 A. Duron, Cumberland Valley — 11 K. Coxe, Palm Beach — 11 J. Boyer, Lancaster — 10 G. Seese, Johnstown-Altoona — 10 C. Schiner, Wyoming — 9 C. Galbavy, Idaho/Montana — 8 C. Gilbertson, Northern Plains — 8 R. Vann, South Carolina — 8 J. Dawson, Pittsburgh — 7 R. Udy, Utah — 7 A. Badeaux, Washington, D.C. — 6 T. Buckler, Columbus — 6 R. Fuller, Green & White Mountains — 6 T. Shirk, Tidewater — 6 K. Temme, Philadelphia — 6 P. Host, Chicago — 5 R. Ledford, Birmingham — 5 P. Strother, New Orleans — 5 W. Wilson, New Orleans — 5 C. Chifici, New Orleans — 4 L. Clark, Milwaukee — 4 J. Ginther, International — 4 C. Griffin, Tulsa — 4 J. Johnson, Northern Plains — 4 J. Reed, Ozark — 4 G. Rolla, L.A./Inland Empire — 4 E. Shreve, Pittsburgh — 4 G. Siepert, Kansas — 4 P. Strother, New Orleans — 4 T. Sumerix, Dayton — 4 R. Zadroga, Philadelphia — 4
SECTIONNEWS
Shown at the Connecticut Section event are Treasurer Walter Chojnacki (left) and Welding Instructor Joseph Hanlon.
Shown at the Central Mass./R.I. Section vendors’ night event are from left (front row) Chair Paul Mendez, Dist rict 1 Direct or Tom Ferri, Douglas Desroc hers, and Robert Winschel, (back row) Brendon Pequita, and Tim Kinnaman.
Shown at the Green & White Section’s tour of Structal-Bridges are (from left) Sherry Morin, Jennifer Eastley, Ann Thompson, Richard Mann, Gerry Ouelette, Geoff Putnam, District 1 Director Tom Ferri, Ernie Plumb, Rich Fuller, tour guide Daryl Hastings, Gary Buckley, Phil Witteman, Feona Lund, John Steel, Perley Lund, and Chris Young.
District 1 Thomas Ferri, director
(508) 527-1884
[email protected]
CENTRAL MASS./R.I. FEBRUARY 28 Activity: The Sect ion held its four th annual welders’ and vendors’ night event at Greater New Bedford Vo-Tech High School. Featured were hands-on demonstrations of new welding and cutting equipment and presentations by local steel and welding-supply companies. The presenters included New Bedford Vo-Tech welding instructors Paul Mendez and Brendon Pe-
quita, Section chair and vice chair, respectively. District 1 Director Tom Ferri par-
lon , welding department head, and Tom Ferri, District 1 director.
ticipated in the event.
CONNECTICUT FEBRUARY 26 Activity: The Section held a students’ night program at Bristol Technical High School in Bristol, Conn., for 65 attendees. The school’s culinary department provided the meals. Attending were representatives from local businesses who discussed their products and services and answered the students’ questions. Assisting with the event were Section Chair Steve Goodrow , Treasurer Walter Chojnacki, Joseph Han-
GREEN & WHITE MTS. FEBRUARY 14 Activity: The Section members toured the Structal-Bridges facility in Claremont, N.H., to study its operations. They ob1 2-in.served submerged arc welding of 2 ⁄ thick flanges to the web of 8- × 100-ft bridge beams. They also toured the welder training area and an environmentally controlled building for painting and thermal spraying. The guide was Daryl Hastings, production trainer. WELDING JOURNAL
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District 2
Harland W. Thompson, director (631) 546-2903
[email protected]
LONG ISLAND FEBRUARY 21 Speaker: Tom Gartland, Section vice chair Affiliation: Trilogy Lab, LLC Topic: Welding Kooks headers Activity: Th e p rogr am was held in Wantagh, N.Y. District 2 Director Harland Thompson attended the program.
PHILADELPHIA
Shown at the Long Island Section program are from left (front row) Tom Gartland, Chair Brian Cassidy, and Alex Duschere, (back row) Barry McQuillan, Ray O’Leary, Jessie Provler, District 2 Director Harland Thompson, and Ken Messemert.
Shown at the March Philadelphia Section program are (from left) District 2 Director Har land Thompson, with Dave Schaffer, Bert Riendeau, Charlie Minnick, Chair Ken Temme, and Tim Stott.
FEBRUARY 7 Speaker: Frank Hauser Affiliation: Divers Academy International Topic: Techniques for testing u nderwater welds Activity: Me mbers of the local chapter of ASNT attended this program. Dist rict 2 Director Harland Thompson and Chair Ken Temme presented the District Director Certificate Award to Mike Chomin, immediate past chair. The program was held at the Crown Plaza in Trevose, Pa. M ARCH 13 Speakers: Dave Schaffer, Airgas East; Bert Riendeau, Airgas Northeast; Charlie Minnick , Miller Electric; and Tim Stott, Miller Electric Topics: Schaffer discussed AC balance control; Riendeau spoke on welding codes made friendly; Minnick talked about traditional welding machine technology; and Stott detailed advances in modern inverter machines Activity: Philadelphia Section Chair Ken Temme presented the District CWI of the Year Award to Bill Mowbray of Scheck Mechanical Contractors; and Section Educator of the Year Awards to Walt Emerle, training director, Plumbers and Pipefitters Local 322; Dan Roskiewich, Gloucester County Institute of Technology; and Charlie Minnick from Miller Electric. The event was hosted by Plumbers and Pipefitters Local 322, a Section supporter, represented by Jim Kehoe, business manager.
District 3
Michael Wiswesser, director (610) 820-9551
[email protected]
Mike Chomin (left) receives the District Di rector Award from District 2 Director Har land Thompson at the February Philadel phia Section event. 72
MAY 2013
Shown at the February Philadelphia Section February m eeting are (from left ) District 2 Director Harland Thompson, Chair Ken Temme, and ASNT Vice Chair Tony Gatti.
District 4
Stewart A. Harris, director (919) 824-0520
[email protected]
Philadelphia Section members display their awards (from left) Dan Roskiewich, Charlie Minnick, Bill Mowbray, Chair Ken Temme, Walt Emerle, and Harland Thompson, District 2 director.
Frank Hauser (left) receives a speak er gift from Ken Temme, Philadelphia Section chair, at the February program.
Tidewater Section members are shown during their tour of Catalina Cylinders in February.
Attendees are shown at the ASME-Triangle Section career panel program in February.
CHARLOTTE CALENDAR M AY 3 13th Annual Intercollegiate Welding Competition Central Piedmont C. C., Charlotte, N.C. Contact Chair Ray Sosko (704) 330-4487
TIDEWATER FEBRUARY 21 Activity: The Sect ion members visited Catalina Cylinders — East, Cliff Impact
Division, Hampton, Va. Following dinner, hosted by the company, Joe Wolff guided the 22 members on a plant tour. The company produces high- and low-pressure aluminum compressed gas cylinders.
M ARCH 14 Speaker: Charlie Pennington Affiliation: ITW Welding North America Topic: Manufacturing filler metals Activity: This Tidewater Section program was held at Smoke BBQ in Newport News, Va., for 28 attendees.
TRIANGLE FEBRUARY 5 Activity: A car eer panel consis ting of six senior members of the local chapter of ASME and Russell Wahrman from the AWS Triangle Section convened at North Carolina University in Raleigh, N.C., to share their experiences in the industry and answer questions for the college students. District 4 Director and Section Chair Stuart Harris expressed an interest in making the career panel program with ASME an annual event.
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Florida West Coast golfers get ready to tee off at the Section’s 21st annual tournament.
Participants are shown at the North Central Florida Section-sponsored welding contest.
Working the Florida West Coast Section raffle-ticket fund-raiser were (from left) Walt Arnold, Bill Machnovitz, Raymond Monson, and Al Sedory.
The top-scoring welders, student category, in the North Central Florida Section contest were Jacob Under hill (right) and Tommy Taylor.
First-place team members at the Florida West Coast tournament are (from left) Walt Arnold, Jack Garrison, Don Chadwell, and Mike Gates.
District 5 Carl Matricardi, director
(770) 979-6344
[email protected]
Joey Pyles and Chair Jenni fer Skyle s took top honors in the professional category at the North Central Florida Section competition. 74
MAY 2013
Dave Parker (left) is shown with speak er Gerry Crawmer at the Northern New York Section meeting.
FLORIDA WEST COAST
NORTH CENTRAL FLORIDA
M ARCH 2 Activity: Th e S ection hosted its 21st annual golf outing to raise funds for its scholarship program to assist local welding students. More than 50 members, guests, and sponsors participated in the event. The first-place team members, scoring 14 under par, were Walt Arnold, Jack Garrison, Don Chadwell, and Mike Gates.
FEBRUARY 12 Activity: The Section conducted a welding contest at Community Technical Adult Education Center in Ocala, Fla., for 11 student and 8 professional welders. Joey Pyles from Townley Industries won the professional title with runner-up Jennifer Skyles, Section chair, with SPX Industries. Jacob Underhill from Bradford Union Area Ca-
Pittsburgh Section and ASNT members are shown at the March program.
Shown at the Tri-State Boy Scout merit badge training session are (kneeling) Allen Black and (from left) Calvin Roach, Xristopher Popoff, Devin Ames, John Saunders, David Hay, Cody Finley, Andy Hall, Sean McKinley, Fred Hammers, and Chad Bowen.
reer Technical Center earned the top student welder honor with Tommy Taylor taking second place. Sebastian Rodriguez from Tech Simulation showed attendees how to test their skills using a virtual arc welding training system.
District 6 Kenneth Phy, director (315) 218-5297
[email protected]
NORTHERN NEW YORK M ARCH 5 Speaker: Gerry Crawmer, welding engineer Affiliation: General Electric (ret.) Topic: Welding steam turbine rotors Activity: The event was held at Shaker Ridge Country Club in Latham, N.Y.
District 7
Uwe Aschemeier, director (786) 473-9540
[email protected]
COLUMBUS M ARCH 19 Speaker: Angie Rybalt, outreach manager Affiliation: American Electric Power (AEP) Topic: How AEP support services can help businesses reduce power consumption Activity: The program was held at La Scala Restaurant in Columbus, Ohio.
Shown at the Pittsburgh Section program are (from left) Chair John Menhart, speaker Wes Williams, and Robert Saunders, ASNT Chapter chair.
DAYTON
TRI-STATE
FEBRUARY 12 Activity: The Section conducted a Boy Scout welding merit badge workshop at Miami Valley Career Technology Center in Clayton, Ohio, for 66 Scouts. Career Center students assisted the Section members to instruct the Scouts on the proper welding techniques. Chuck Ford, Section student affairs chair, and Chair Chris Lander led the activity.
J ANUARY 8, 10, 15 Activity: Fred Hammers, owner of Hammers Industries, and Health and Safety Manager Cody Finley hosted and trained a Boy Scouts welding merit badge class. The hands-on shop training was taught by Shop Manager Calvin Roach, Andy Hall, and John Saunders . Earning their welding badges were David Hay, Xristopher Popoff , and Chad Bowen of Troop 63, Devin Ames of Troop 790, Sean McKinley of Troop 50, and Allen Black of Troop 92. The training was conducted at the Hammers Industries facilities near Huntington, W.Va.
FEBRUARY 16, 17 Activity: The Dayton Section members participated in Tech Fest 2013 presented by the Affiliated Societies of Dayton to interest grade school students in science and technology. The Section members talked to students about careers in welding and presented live demonstrations of automated welding using a Motoman robot.
PITTSBURGH M ARCH 12 Speaker: Wes Williams, staff engineer Affiliation: First Energy Corp. Topic: Procedure for repair of cracks in the Unit 2 Nuclear Reactor head Activity: Members of the local chapter of ASNT International Chapter, headed by Chair Robert Saunders, attended this program, held at Springfield Grille in Mars, Pa.
District 8
Joe Livesay, director (931) 484-7502, ext. 143
[email protected]
CHATTANOOGA FEBRUARY 16 Activity: The Section held its student welding competition at Sequoyah High School in Soddy Daisy, Tenn., hosted by the Sequoyah High School Student Chapter. The top scorers in the high school category were Dustin Luthringer, Colton Jones, and Dexter McSpaden. The postsecondary category winners were Mauricio Gayton Jr., Eric Vennie, and Darren Vincent. WELDING JOURNAL
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HOLSTON VALLEY M ARCH 12 Speaker: Gary Roberts Affiliation: Airgas, automation specialist Topic: Airgas University online training Activity: Nominations for Section officers were received for presentation at the April meeting. The program was held at Mama’s House Restaurant in Kingsport, Tenn. Shown are the participants in the Chattanooga Section/Sequoyah High School Student Chapter welding contest.
The top-scoring welders in the Chattanooga Section/Sequoyah High School Student Chapter contest are (from left) Eric Vennie, Dexter McSpaden, Dustin Luthringer, Colton Jones, Darren Vincent, and Mauricio Gayton Jr.
NORTHEAST TENNESSEE FEBRUARY 12 Activity: The Section hosted its annual students’ day event at Tennessee Technology Center (TTC) in Knoxville for about 230 students from local schools. The event included talks on welding careers, complimentary lunch, welding demonstrations, and a tour of the Technology Center. Attending were high school welding instructors Steve Linn (TTC Knoxville), Jeff Hankins (Oak Ridge), Jim Thomas (South Doyle), Rick Johnson (Grainger), Tim Steelman (Morgan County), Michael Hurt (TTC Jacksboro), Mike Russell (TTC Harriman), and District 8 Director Joe Livesay (TTC Crossville).
NORTHEAST MISSISSIPPI FEBRUARY 21 Activity: The Section sponsored a students’ night and equipment demo event featuring the Lincoln Electric mobile display unit. Lincoln salesman Ron Martucci demonstrated the equipment. Speaker Gary Roberts (red sweater) is shown with the Holston Valley Section members.
District 9
George Fairbanks Jr., director (225) 473-6362 fi
[email protected]
BIRMINGHAM
Students are shown at the Northeast Tennessee Section’s students’ day event.
FEBRUARY 27 Activity: The Birmingham Section and Lawson State C. C. Student Chapter members held a welding seminar at Plumbers and Pipefitters Local 372 in Duncanville, Ala. Exhibitors included Victor, Lincoln Electric, Miller E lectric, Airgas Welding Supply, and Harris Equipment.
Central Alabama Student Chapter 7 Speaker: Craig Ray Topic: Jobs in underwater welding and commercial diving Activity: The program was held in the Central Alabama Community College welding shop in Alexander, Ala. Attending were D. MARCH
Steve Linn answers students’ questions at the Northeast Tennessee event. 76
MAY 2013
J. James, Emily Hatfield, Daniel Arnberg, Brandon Fraser, Robin Holt, Zack Adams, Chris Floyd, Walter Whatley, Colton Stroud, and Craig Ray.
Ron Martucci is shown at the Northeast Mis sissippi Section students program. Northeast Mississippi Section members and students are shown at the February event.
NEW ORLEANS FEBRUARY 26 Speaker: Nancy Cole, AWS president Affiliation: NCC Engineering Topic: The welding industry Activity: Past AWS Pres iden t John and wife Donna Bruskotter hosted a reception for Nancy Cole at their home in Slidell, La. District 9 Director George Fairbanks was among the Section’s guests who attended the event. M ARCH 19 Speaker: Jason Lange Affiliation: Lincoln Automation, Inc. Topic: Controlling welding fumes Activity: IWS Gas & Su pply, represented by President Moussy Chassion , provided the door prizes and sponsored this New Orleans Section program at Café Hope in Marrero, La., for 66 members, students, and guests. Alfred Ma rshal l , an apprentice with Ironworkers No. 58, was the 50/50 prize winner.
Attendees are shown at the Birmingham Section-sponsored welding seminar.
District 10
Robert E. Brenner, director
(330) 484-3650
[email protected]
District 10 M ARCH 9 Activity: The District held its second CWI Roundtable event at Babcock & Wilcox Commercial in Euclid, Ohio. The event offers CWIs an opportunity to share their experiences and opinions. The 20 participants discussed a number of topics, including qualification records, welding procedure specifications, welder performance qualification records, and the proper way to issue welder certification papers.
Central Alabama Student Chapter members are shown at their March meeting.
DRAKE WELL J ANUARY 18 Activity: The S ection p articipated in the SkillsUSA competition hosted by the New Castle School of Trades in New Castle, Pa. The top welders were Joseph Crate, Robert Ackerson, and Conner Biggs.
Ironworker Alfred Marshall (left) receives a door pri ze from Aldo Duron, New Orleans Section chair.
Nancy Cole, AWS president, is shown with George Fairbanks (left), District 9 director, and D. J. Berger at the New Orleans Section reception. WELDING JOURNAL
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The New Orleans Section attendees are shown at the reception held at the Bruskotter’s home.
District 10 CWIs are shown during their roundtable discussion in March. M ARCH 12 Activity: The Drake Well Section members discussed the March 9 District 10 Roundtable event and the District 10 conference scheduled for April 19. Dan Bubenhiem was elected vice chair.
MAHONING VALLEY
Shown at the New Orleans Section event are (from left) Tommy Callahan, Mike Massicot, Jimmy Gibbs, Moussy Chassion, and Chair Aldo Duron.
M ARCH 14 Speaker: Jim Hannahs, PE, CWI Topic: Metals and processes used to build NASCAR vehicles Activit y: The meeting, h eld at Mahon ing County Career & Technical Center in Canfield, Ohio, was attended by about 90 members and guests.
NORTHWESTERN PENNSYLVANIA FEBRUARY 13 Speaker: Marty Siddall , technical sales representative/automation specialist Affiliation: The Lincoln Electric Co. Topic: New techno logies in robo tic welding Activity: The meeting was held at Tri State Business Institute in Erie, Pa.
District 11 Robert P. Wilcox, director
Shown at the New Orleans Section March program are (from left) Moussy Chassion, Matt Howerton, speaker Jason Lange, and Chair Aldo Duron. 78
MAY 2013
(734) 721-8272
[email protected]
NORTHWESTERN OHIO FEBRUARY 27 SPEAKER: Curt Wilsoncroft, regional sales representative Affiliation: Victor Technologies Topic: Carbon arc and oxygen lance cutting Activity: The program was held at Owens Community College in Perrysburg, Ohio, for 41 attendees. Following the talk, the group visited the welding lab where Wilsoncroft demonstrated the cutting processes and members had a hands-on opportunity to experiment with the equipment.
Shown at the Drake Well Section SkillsUSA competition are (from left) Robert Ackerson, Tyler Hoffman, Joseph Crate, Chad Hajec, Joseph Steiner, and Conner Biggs.
District 12 Daniel J. Roland, director
(715) 735-9341, ext. 6421 daniel.roland@us.fincantieri.com
RACINE-KENOSHA FEBRUARY 22 Speaker: Chris Boycks, CWI Topic: The shortage of skilled qualified welders and fabricators Activity: The seventh annual Dist rict 12 winte r meeting was hoste d by Ja y Manufacturing Co. in Oshkosh, Wis. Dan Roland , District 12 director, presented Chair Dan Crifase the Dalton E. Hamilton Memorial CWI of the Year Award and the District Meritorious Award to Vice Chair Ken Karwowski.
Drake Well Section members are from left (f ront row) Robert Fugate, Colin Young, Justus Burk, Bailey Hagerty, Dan Bubenhiem, William Brownlee, Travis Crate, and Ward Kiser; (back row) Rolf Laemmer, Erick Speer, Joe Crate, and Troy Braden.
District 13 John Willard, director
(815) 954-4838
[email protected]
CHICAGO FEBRUARY 13 Activity: The Section hosted its annu al St. Valentine’s Day dinner party at Cooper’s Hawk Winery & Restaurant in Chicago, Ill., for 35 attendees.
Speaker Jim Hannahs (right) is shown with Dave Hughes, Mahoning Valley Section ex ecutive committee member.
Shown at the February Racine-Kenosha Section event are (from left) Chair Dan Crifase, District 12 Director Dan Roland, and Ken Karwowski.
From left, Gary Dugger, Bennie Flynn, and Gary Tucker worked the Indiana SkillsUSA contest.
Speaker Marty Siddall (left) is shown with Tom Kostreba, Northwestern Pennsylvania Section chair.
District 14
Robert L. Richwine, director
(765) 378-5378
[email protected]
INDIANA FEBRUARY 9, 10 Activity: The Sect ion cond ucte d the regional SkillsUSA welding contest at J. E. L. Career Center in Indianapolis, Ind., for 20 participants. Four secondary and four postsecondary welding contestants were selected to compete in the state welding contest. Working the event were Bob Rich wine , District 14 director; Chair Bennie Flynn; Gary Dugger; and Gary Tucker.
WELDING JOURNAL
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Shown at the Lexington Section program are (from left) Rosa Whitaker, Coy Hall, Rosa McCallum, and Shawn Gannon.
SkillsUSA welding contestants are shown during the Indiana Section event.
FEBRUARY 27 Activity: The Indiana Section members toured the Don Schumacher Racing facility in Brownsburg, Ind. Mike Lewis, senior vice president, conducted the tour.
LEXINGTON
Shown at the Indiana Section contest are (from left) Wilson Smith, Thomas Faucett, and Bob Richwine, District 14 director.
Incoming Lexington Section Chair Coy Hall (left) is shown with Sherman Cook.
FEBRUARY 28 Speaker: Tony Noah Affiliation: The Lincoln Electric Co. Topic: Pulse welding Activity: Incoming Chair Coy Hall presented the Section CWI of the Year Award to Sherman Cook , an instructor at Rockcastle County Technical School. Rose Whitaker and Rosa McCallum were presented $250 scholarship awards. Fifty members and guests attended the program.
ST. LOUIS DECEMBER 21 Activity: The Section held its holiday party at Cee Kay Supply, Inc., in St. Louis, Mo. Company owner Tom Dunn was cited f or his generous support and the services he has offered to the Section over the years. In recognition, Jerry Simpson presented Dunn the District 14 Meritorious Award.
Angela Harrison is shown with Dennis Pick ering at the Arkansas Welding Expo.
Mike Lewis led the Indiana Section mem bers on a tour of Don Schumacher Racing.
District 15 David Lynnes, director (701) 365-0606
[email protected]
District 16 Dennis Wright, director (913) 782-0635
[email protected]
Shown at the February Central Arkansas Section program are (from left) Aaron Carr, Karen Cooper, and Dennis Pickering. 80
MAY 2013
Tom Dunn (left) receives the District 14 Meritorious Award from Jerry Simpson at the St. Louis Section holiday party.
District 17
J. Jones, director (832) 506-5986
[email protected]
Speaker Dennis Pickering (kneeling) and attendees are shown at the February Central Arkansas Section meeting.
CENTRAL ARKANSAS NOVEMBER 1 Activity: The Section participated in the Arkansas Welding Expo presented by WELSCO, Inc., at Verizon Arena in Little Rock, Ark. Angela Harrison, WELSCO president, was presented the Section Meritorious Award. More than 700 students attended the event to learn about job opportunities from Monica Pfarr, AWS corporate director, workforce development, who held presentations for them titled “Let the Sparks Fly.” Vice Chair Dennis Pickering, a welding instructor at Arkansas Career Training Institute, made presentations at the expo.
Attendees are shown at the Central Arkansas Section January program.
J ANUARY 17 Speaker: Dennis Pickering, vice chairman Affiliation: Arkansas Career Training Institute, welding instructor Topic: Welding codes and standards Activity: The program was held at the institute in Hot Springs, Ark. FEBRUARY 11 Speaker: Dennis Pickering, vice chairman Affiliation: Arkansas Career Training Institute, welding instructor Topic: The AWS scholarship program Activity: The meeting was held at Arkansas State University in Heber Springs, Ark.
Shown at the East Texas Section program are (from left) Student Chapter Chair Michael Florczykowski, Chair Bryan Baker, Yoni Adonyi, speaker Tom Siewert, Robert Warke, and J. Jones, District 17 director.
EAST TEXAS LeTourneau University S. C. FEBRUARY 21 Speaker: Tom Siewe rt, AWS director-atlarge Affiliation: NIST (ret.) Topic: Analysis of the collapse of the World Trade Center buildings Activity: The program was held at LeTourneau University in Longview, Tex. Attending were District 17 Director J. Jones, Chair Bryan Baker, Student Chapter Chair Michael Florczykowski, Prof. Materials Joining Engineering Yoni Adonyi , and Robert Warke, associate professor, materials joining.
Shown at the Tulsa Section program are (from left) AWS President Nancy Cole, Todd Mor ris, Charles Griffin, Ralph Johnson, Ray Wilsdorf, and J. Jones, District 17 director.
TULSA FEBRUARY 9 Activity: The Section hosted a dinner to celebrate St. Valentine’s day with the ladies. Nancy Cole, AWS president, attended the event. Todd Morris and Ralph
Johnson received District Director Certificate Awards, and Charles Griffin was pre-
sented the Private Sector Instructor Award. Ralph Johnson and Ray Wilsdorf received Dalton E. Hamilton Memorial CWI of the Year Awards. WELDING JOURNAL
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Attendees are shown at the Houston Section instructors institute held February 22.
Shown at the Houston Section booth at the Houstex® show are (from left) Sam Gentry, Luanne Bray, John Stoll, and John Bray, District 18 director.
Tac Edwards (left), Lake Charles Section chair, presents a speak er plaqu e to John Bray, District 18 director.
Alaska Section members and guests are shown at the February program.
Members and students are shown at the Spokane Section program in January.
District 18 Shown during the British Columbia Section tour are Brad Moe and speaker John Shaw. 82
MAY 2013
John Bray, director (281) 997-7273 sales@affiliatedmachinery.com
HOUSTON FEBRUARY 22 Activity: The Section held its seven th annual instructors institute hosted by Andre Horn and the Industrial Welding Academy staff. Thirty welding instructors attended
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Idaho-Montana Section members are shown during their tour of Spudnik Equipment Co.
Dale Flood (left), an AWS director-at-large, accepts a speaker gift from Brad Bosworth, Fresno Section chair. the event. The topic for the hands-on class was flux cored arc we lding. Each in structor had the opportunity to qualify on a sin3 gle V-groove butt joint on ⁄ 8-in. plate in either the flat or vertical position. The steel plates were donated by Scott Witkowski and Maverick Testing Laboratories. FEBRUARY 26–28 Activity: The Se ction manned a booth at the Houstex® The Art of Manufacturing Show presented by the Society of Manufacturing Engineers at George R. Brown Convention Center in Houston, Tex. Participating were Sam Gentry, executive director, AWS Foundation; Joe Krall, AWS managing director, global exposition sales; John Stol l, and Luanne and John Bray , District 18 director.
LAKE CHARLES FEBRUARY 20 Speaker: John Bray, District 18 director Affiliation: Affiliated Machinery Topic: What’s new at AWS Activity: The dinn er and program were held at L ogan’s Roadhouse Restaurant in Lake Charles, La.
Attendees are shown at the Fresno Section program in February.
BRITISH COLUMBIA FEBRUARY 20 Speaker: John Shaw , VP, government relations and business development Affiliation: Seaspan Vancouver Shipyards Topic: Update on the National Shipbuilding Procurement Strategy Activity: Following a catered din ner and the talk, the Section members were guided on a tour of the Seaspan Vancouver Ship yards by AWS member Brad Moe.
SPOKANE J ANUARY 16 Speaker: Russ Loveland Affiliation: Western States Equipment Co. Topic: Maintaining large construction equipment Activity: The program was held at Spokane Community College in Spokane, Wash., for 39 attendees.
District 20
William A. Komlos, director (801) 560-2353 [email protected]
District 22 Kerry E. Shatell, director (925) 866-5434 [email protected]
FRESNO FEBRUARY 21 Activity: Th e Section membe rs met for a demonstration of the Tri Tool AdaptArc orbital welding system. Dale Flood, Tri Tool project manager, an AWS directorat-large, and past District 22 director, conducted the program for about 45 attendees. Attending were District 22 Director Kerry Shatell , Kent Baucher, a past District 22 director, and Theo Davis , an instructor at Fresno City College.
SACRAMENTO VALLEY J ANUARY 16 Speaker: Mark Paavola , administrator of apprenticeship ad training Affiliation: Sheet Metal Workers Assn. Topic: Employment opportunities in the sheet metal trade Activity: Paavola and David Perez detailed the training program used by the center.
IDAHO-MONTANA
District 19
Ken Johnson, director (425) 957-3553 [email protected]
ALASKA FEBRUARY 20 Speaker: Marty Anderson , chair, ASNT Alaska chapter Affiliation: Alaska Technical Training, Inc. Topic: Nondestructive welding inspection technologies Activity: The progr am was held for 21 attendees in Anchorage, Alaska.
FEBRUARY 13 Activity: The Sect ion memb ers visited Spudnik Equipment Co., in Blackfoot, Idaho, to tour the facility. The facility designs and manufactures potato planting, cultivating, harvesting, and handling equipment. Wes Woodland , shift supervisor, led the program.
District 21
Nanette Samanich, director (702) 429-5017 [email protected]
International GERMANY C ALENDAR
Essen, Germany SEPT. 11–17 66th IIW Annual Assembly 2013 Int’l Trade Fair Joining, Cutting, Surfacing
SEPT. 16, 17 Int’l Conf. on Automation in Welding
SEPT. 16–21 Young Welders’ Competitions
www.iiw2013.com WELDING JOURNAL
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Guide to AWS Services
American Welding Society 8669 Doral Blvd., Ste. 130, Doral, FL 33166 (800/305) 443-9353; FAX (305) 443-7559; www.aws.org Staff phone extensions are shown in parentheses. AWS PRESIDENT
Nancy C. Cole [email protected]
NCC Engineering 2735 Robert Oliver Ave. Fernandina Beach, FL 32034
ADMINISTRATION Executive Director Ray W. Shook . [email protected] . . . . . . . . . .(210)
INTERNATIONAL SALES Managing Director, Global Exposition Sales Joe Krall.. [email protected] . . . . . . . . . . . . . . . .(297) Corporate Director, International Sales
Jeff P. Kamentz .. [email protected] . . . . . . .(233) Oversees international business activities involving cert ificat ion, p ublica tion, and me mbershi p.
TECHNICAL SERVICES Department Information . . . . . . . . . . . . . . . . .(340) Managing Director Andrew R. Davis.. [email protected] . . . . . . .(466) International Standards Activities, American Council of the International Institute of Welding (IIW) Director, National Standards Activities
Annette Alonso.. [email protected] . . . . . . .(299)
.
Sr. Associate Executive Director Cassie R. Burrell.. [email protected] . . . . . .(253)
PUBLICATION SERVICES Department Information . . . . . . . . . . . . . . . . .(275) Managing Director Andrew Cullison . [email protected] . . . . . .(249) .
Chief Financial Officer
Gesana Villegas.. [email protected] . . . . . .(252)
Welding Journal
VP Sales and Marketing Bill Fudale.. [email protected] . . . . . . . . . . . . .(211)
Andrew Cullison . [email protected] . . . . . .(249)
VP Technology and Business Development
Editor Mary Ruth Johnsen . [email protected] . .(238)
Executive Assistant for Board Services Gricelda Manalich.. [email protected] . . . . .(294)
Rob Saltzstein.. [email protected] . . . . . . . . . . .(243)
Dennis Harwig.. [email protected] . . . . . . . . .(213)
Administrative Services Managing Director Jim Lankford.. [email protected] . . . . . . . . . . . . .(214) IT Network Director
Armando Campana.. [email protected] . .(296)
Director
Hidail Nuñez. [email protected] . . . . . . . . . . . .(287)
Director of IT Operations
Natalia Swain.. [email protected] . . . . . . . . . .(245)
Human Resources Director, Compensation and Benefits Luisa Hernandez.. [email protected] . . . . . . . . .(266)
Publisher
.
.
National Sales Director
Society and Section News Editor Howard Woodward [email protected] . .(244) ..
Welding Handbook
RWMA — Resistance Welding Manufacturing Alliance Management Specialist Keila DeMoraes.... [email protected] . . . .(444) WEMCO — Association of Welding Manufacturers Management Specialist Keila DeMoraes.... [email protected] . . . .(444) Brazing and Soldering Manufacturers’ Committee Jeff Weber.. [email protected] . . . . . . . . . . . . .(246) GAWDA — Gases and Welding Distributors Association Executive Director John Ospina.. [email protected] . . . . . . . . . .(462) Operations Manager
Natasha Alexis.. [email protected] . . . . . . . . .(401)
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Senior Staff Engineer
Rakesh Gupta.. [email protected] . . . . . . . . . .(301)
Filler Metals and Allied Materials, International Filler Metals, UNS Numbers Assignment, Arc Welding and Cutting Processes Standards Program Managers
Stephen Borrero... [email protected] . . . . .(334)
Public Relations Manager
Cindy Weihl.. [email protected] . . . . . . . . . . . .(416)
Webmaster Jose Salgado.. [email protected] . . . . . . . . .(456)
MEMBER SERVICES Department Information . . . . . . . . . . . . . . . . .(480) Sr. Associate Executive Director Cassie R. Burrell.. [email protected] . . . . . .(253)
ITSA — International Thermal Spray Association Senior Manager and Editor Kathy Dusa. [email protected] . . .(232)
Structural Welding, Methods of Inspection, Mechanical Testing of Welds, Welding in Marine Construction, Piping and Tubing
MARKETING COMMUNICATIONS Director Ross Hancock.. [email protected] . . . . . . .(226)
International Institute of Welding Senior Coordinator Sissibeth Lopez . . [email protected] . . . . . . . . .(319) Liaison services with other national and international societies and standards organizations.
CONVENTION and EXPOSITIONS Director, Convention and Meeting Services Matthew Rubin..... [email protected] . . . . . . .(239)
Managing Engineer, Standards
Brian McGrath .... [email protected] . . . . .(311)
Efram Abrams.. [email protected] . . . . . . . .(307)
Section Web Editor Henry Chinea... [email protected] . . . . . . . . .(452)
GOVERNMENT LIAISON SERVICES
Metric Practice, Safety and Health, Joining of Plastics and Composites, Welding Iron Castings, Personnel and Facilities Qualification
Editor Annette O’Brien.. [email protected] . . . . . . .(303)
Director, Human Resources Dora A. Shade.. [email protected] . . . . . . . . .(235)
Hugh K. Webster . . . . . . . . . [email protected] Webster, Chamberlain & Bean, Washington, D.C., (202) 785-9500; FAX (202) 835-0243. Monitors fed eral issues of importance to the industry.
Manager, Safety and Health
Stephen P. Hedrick .. [email protected] . . . . . .(305)
Director
Rhenda A. Kenny... [email protected] . . . . . .(260) Serves as a liaison between Section members and AWS headquarters.
CERTIFICATION SERVICES Department Information . . . . . . . . . . . . . . . . .(273) Managing Director John L. Gayler.. [email protected] . . . . . . . . . .(472) Oversees all certification activities including all inter national certification programs. Director, Certification Operations
Terry [email protected] . . . . . . . . . . . . .(470) Oversees application processing, renewals, and exam scoring.
Director, Certification Programs
Linda Henderson.. [email protected] . . . . . . .(298) Oversees the development of new certification pro grams, as well as AWS-Accredited Test Facilities, and AWS Certified Welding Fabricators.
EDUCATION SERVICES Director, Operations Martica Ventura.. [email protected] . . . . . .(224) Director, Education Development
David Hernandez.. [email protected] . . .(219)
AWS AWARDS, FELLOWS, COUNSELORS Senior Manager Wendy S. Reeve.. [email protected] . . . . . . . .(293)
Coordinates AWS awards, Fellow, Counselor nominees.
Thermal Spray, Automotive, Resistance Welding, Machinery and Equipment Brazing and Soldering, Brazing Filler Metals and Fluxes, Brazing Handbook, Soldering Handbook, Railroad Welding, Definitions and Symbols Alex Diaz.... [email protected] . . . . . . . . . . . . . .(304)
Welding Qualification, Sheet Metal Welding, Aircraft and Aerospace, Joining of Metals and Alloys Patrick Henry.. [email protected] . . . . . . . . . .(215)
Friction Welding, Oxyfuel Gas Welding and Cutting, High-Energy Beam Welding, Robotics Welding, Welding in Sanitary Applications Senior Manager, Technical Publications
Rosalinda O’Neill.. [email protected] . . . . . . .(451)
AWS publish es about 2 00 docum ents wide ly used throughout the welding industry
Note : Officia l int erpr etations o f AWS s tanda rds may be obtained only by sending a request in writing to Andrew R. Davis, managing director, Tech nical Services, [email protected]. Oral opinions on AWS standards may be ren dered, however, oral opinions do not constitute of fici al or unoff icia l opi nions o r inte rpretations of AWS. In addition, oral opinions are informal and shou ld not be use d as a s ubsti tute f or an offic ial interpretation. AWS FOUNDATION, Inc.
www.aws.org/w/a/foundation General Information (800/305) 443-9353, ext. 212, [email protected]
Chairman, Board of Trustees Gerald D. Uttrachi
Executive Director, Foundation
Sam Gentry.. [email protected] . . . . . . . . . . . . . . . (331)
Corporate Director, Workforce Development
Monica Pfarr.. [email protected]. . . . . . . . . . . . . . . . (461)
The AWS Foundation is a not-for-profit corporation established to provide support for the educational and scientific endeavors of the American Welding Society. Promote the Foundation’s work with your financial support. Call (800) 443-9353, ext. 212, for complete information.
PERSONNEL
Eriez® Names Two to Key Positions Eriez®, Erie, Pa., a supplier of magnetic lift, conveying, metal-detection, X-ray, controlling, and inspection equipment, has promoted Andrew Goldner to
senior manager exports, and John Blicha to director of corporate communications. Goldner will manage the company’s sales offices in Central and South America and the Middle East. Most recently, Goldner served as export market development manager, targeting industrial markets in Central and South America. Blicha has Andrew Goldner
John Blicha
served as marketing communications manager since joining the company last year.
Tuchscherer Elected AWS Foundation Trustee
2% Thoriated
E3® Electrodes were te ste d on w hen compared to
after 3 passes
The American Welding Society Foundation, Inc., Doral, Fla., has elected Becky Tuchscherer to serve on its board of trustees. Her term runs through 2015. Tuchscherer is group vice president, commercial welding, for Miller Electric Mfg. Co., where she has worked since 1988. She is an AWS member and has served on the AWS Finance Committee. The AWS Foundation was established in 1991 to support programs that ensure the growth and development of the welding industry. Its focus is on providing scholarships for welding students and pursuing welder workforce de velopment issues. Becky Tuchscherer
Wall Colmonoy Announces New Controller and CFO Wall Colmonoy Corp., Madison Heights, Mich., a supplier of surfacing and brazing products, castings, and engineered components, has named Michael Edwards chief financial officer, and Michael Safford controller for the company’s U.S. operations. Edwards has ex-
E3® after 8 passes
E3®
For info go to www.aws.org/ad-index
Mike Edwards
Mike Safford — continued on page 90
88
MAY 2013
— continued from page 88
tensive experience in accounting with Deloitte and in manufacturing where he has held senior finance and operations management positions with a machine manufacturer and a provider of injectionmolded PVC pipe fittings. Safford is a certified managerial accountant, with Six Sigma Green Belt training.
Centerline® Names RW Account Manager Centerline® (Windsor) Ltd., Windsor, Ont., Canada, has named Greg Van Dyke account manager specializing in resistance welding consumables and automation component products. Van Dyke has ten years of experience in the resistance welding field, most recently with Resistance Welding Products Ltd., and the Tuffaloy Group of companies. Greg Van Dyke
Sales Director Hired at Caster Concepts
Adept Technology Appoints Global Marketing VP
Caster Concepts, Albion, Mich., a supplier of heavy-duty casters and wheels, has named Jami e Long director of sales. Prior to joining the company, Long served as national sales manager for Hotsy Corp., a division of Karcher North America, a manufacturer of high-pressure clean Jamie Long ing systems.
Adept Technology, Inc., Pleasanton, Calif., a provider of intelligent robots, has named Glenn Hewson senior vice president of global marketing. Hewson, with 25 years of product line and marketGlenn Hewson ing experience, most recently served as vice president of global marketing at Avure Technologies, Inc.
TRUMPF Designates West Coast Sales Manager TRUMPF, Inc., Laser Technology Center, Farmington, Conn., has named Gene Bonacum west coast regional sales manager to handle accounts in Oregon, Washington, California, Idaho, Montana, and Wyoming. Bonacum has more than 20 years of experience in the field, most recently with Newport Corp.
For info go to www.aws.org/ad-index
For info go to www.aws.org/ad-index
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MAY 2013
Member Milestones Damian J. Kotecki Damian J. Kotecki, an AWS Fellow, PE, a past AWS president, and worldrenowned authority on welding stainless steels, has been appointed to the Board of Advisors of Abakan, Inc., Miami, Fla. The announcement states in part, “Dr. Kotecki brings 43 years of welding expertise as well as his extensive technical and business net Damian Kotecki work to the Board of Advisors, as Abakan transitions its CermaClad™ large-area cladding technology into full commercial production. Dr. Kotecki’s extensive experience in welding research, pipeline failure analyses, welding training and specifications, welding procedure development, quality assurance, and stainless/high-alloy welding filler metal and product development will help assure the company’s products incorporate the highest levels of technical excellence.” Kotecki chairs the A5D Subcommittee on Stainless Steel Welding and the International Standards Activities Committee. He is also a past chair of the IIW Commission II Arc Welding and Filler Metals, and authors the bimonthly Stainless Q&A column in the Welding Journal . He conducted welding research projects and pipeline failure analyses for the Battelle Memorial Institute, served as director of research for Teledyne McKay, and most recently retired as the technical director for stainless and high-alloy product development for The Lincoln Electric Co. ◆
awo.aws.org
Understanding Under standing Welding W elding Symbols Knowledge of weld joint terminology is essential for all levels of the welding design and production process. Use of proper terms makes it much easier for welding personnel to communicate about various fit-up and welding problems encountered during the fabrication process. A A welding welding inspector inspect inspector’s ’s ability to read and and interpret interpret welding welding plans correctly is essential to properly inspecting a piece or part. This in-depth course walks the user through A AWS WS 2.4:2012, 2.4:2012, starting starting with with a module on orthographic orthographic views, views, joint joint types, and weld types. Then the course dives into the t he various various types types of welds and clarifies the rules and usage usage of of welding symbols. This self-paced course covers basic joint geometry geometry,, groove groove welds, welds, fillet fillet welds, welds, plug plug and and slot slot welds, spot and projection welds, and stud, seam, surfacing, and edge welds. Rounding out the seminar is a module on brazing terms and symbols and non-destructive testing symbols. Interactive practice problems include an explanation of each solution, and chapter quizzes will solidify the knowledge and prepare you for the proficiency exam. The seminar is approximately 12 hours long and concludes with a final test.
Sample seminar at awo.aws.org/seminars/symbols
Friends and Colleagues:
The American Welding Society established the honor of Counselor to recognize individual members for a career of distinguished organizational leadership that has enhanced the image and impact of the welding industry. Election as a Counselor shall be based on an individual’s career of outstanding accomplishment. To be eligible for appointment, an individual shall have demonstrated his or her leadership in the welding industry by one or more of the following: • Leadership of or within an organization that has made a substantial contribution to the welding industry. The individual’s organization shall have shown an ongoing commitment to the industry, as evidenced by support of participation of its employees in industry activities. • Leadership of or within an organization that has made a substantial contribution to training and vocational education in the welding industry. The individual’s organization shall have shown an ongoing commitment to the industry, as evidenced by support of participation of its employees in industry activities. For specifics on the nomination requirements, please contact Wendy Sue Reeve at AWS headquarters in Miami, or simply follow the instructions on the Counselor nomination form in this issue of the Welding Journal. The deadline for submission is July 1, 2013. The committee looks forward to receiving these nominations for 2014 consideration.
Sincerely, Lee Kvidahl Chair, Counselor Selection Committee
Nomination of AWS Counselor
I.
HISTORY AND BACKGROUND In 1999, the American Welding Society established the honor of Counselor to recognize individual members for a career of distinguished organizational leadership that has enhanced the image and impact of the welding industry. Election as a Counselor shall be based on an individual’s career of outstanding accomplishment. To be eligible for appointment, an individual shall have demonstrated his or her leadership in the welding industry by one or more of the following: • Leadership of or within an organization that has made a substantial contribution to the welding industry. (The individual’s organization shall have shown an ongoing commitment to the industry, as evidenced by support of participation of its employees in industry activities such as AWS, IIW, WRC, SkillsUSA, NEMA, NSRP SP7 or other similar groups.) • Leadership of or within an organization that has made substantial contribution to training and vocational education in the welding industry. (The individual’s organization shall have shown an ongoing commitment to the industry, as evidenced by support of partici pation of its employees in industry activities such as AWS, IIW, WRC, SkillsUSA, NEMA, NSRP SP7 or other similar groups.) II. RULES A. B. C. D. E. F. G.
Candidates for Counselor shall have at least 10 years of membership in AWS. Each candidate for Counselor shall be nominated by at least five members of the Society. Nominations shall be submitted on the official form available from AWS headquarters. Nominations must be submitted to AWS headquarters no later than July 1 of the year prior to that in which the award is to be presented. Nominations shall remain valid for three years. All information on nominees will be held in strict confidence. Candidates who have been elected as Fellows of AWS shall not be eligible for election as Counselors. Candidates may not be nominated for both of these awards at the same time.
III. NUMBER OF COUNSELORS TO BE SELECTED Maximum of 10 Counselors selected each year. Return completed Counselor nomination package to: Wendy S. Reeve American Welding Society Senior Manager Award Programs and Administrative Support 8669 Doral Blvd., Suite 130 Doral, FL 33166 Telephone: 800-443-9353, extension 293
SUBMISSION DEADLINE: July 1, 201 3
(please type or print in black ink)
CLASS OF 2014 COUNSELOR NOMINATION FORM
DATE_________________NAME OF CANDIDATE________________________________________________________________________ AWS MEMBER NO.___________________________YEARS OF AWS MEMBERSHIP____________________________________________ HOME ADDRESS____________________________________________________________________________________________________ CITY_______________________________________________STATE________ZIP CODE__________PHONE________________________ PRESENT COMPANY/INSTITUTION AFFILIATION_______________________________________________________________________ TITLE/POSITION____________________________________________________________________________________________________ BUSINESS ADDRESS________________________________________________________________________________________________ CITY______________________________________________STATE________ZIP CODE__________PHONE_________________________ ACADEMIC BACKGROUND, AS APPLICABLE: INSTITUTION______________________________________________________________________________________________________ MAJOR & MINOR__________________________________________________________________________________________________ DEGREES OR CERTIFICATES/YEAR____________________________________________________________________________________ LICENSED PROFESSIONAL ENGINEER: YES_________NO__________ STATE______________________________________________ SIGNIFICANT WORK EXPERIENCE: COMPANY/CITY/STATE_____________________________________________________________________________________________ POSITION____________________________________________________________________________YEARS_______________________ COMPANY/CITY/STATE_____________________________________________________________________________________________ POSITION____________________________________________________________________________YEARS_______________________ SUMMARIZE MAJOR CONTRIBUTIONS IN THESE POSITIONS: __________________________________________________________________________________________________________________ __________________________________________________________________________________________________________________ __________________________________________________________________________________________________________________ IT IS MANDATORY THAT A CITATION (50 TO 100 WORDS, USE SEPARATE SHEET) INDICATING WHY THE NOMINEE SHOULD BE SELECTED AS AN AWS COUNSELOR ACCOMPANY THE NOMINATION PACKET. IF NOMINEE IS SELECTED, THIS STAT EMENT MAY BE INCORPORATED WITHIN THE CITATION CERTIFICATE. **MOST IMPORTANT**
The Counselor Selection Committee criteria are strongly based on and extracted from the categories identified below. All information and support material provided by the candidate’s Counselor Proposer, Nominating Members and peers are considered. SUBMITTED BY: PROPOSER_______________________________________________ AWS Member No.___________________ The proposer will serve as the contact if the Selection Committee requires further information. The proposer is encouraged to include a detailed biography of the candidate and letters of recommendation from individuals describing the specific accomplishments of the candidate. Signatures on this nominating form, or supporting letters from each nominator, are required from four AWS members in addition to the proposer. Signatures may be acquired by photocopying the original and transmitting to each nominating member. Once the signatures are secured, the total package should be submitted.
NOMINATING MEMBER:___________________________________Print Name___________________________________ AWS Member No.______________ NOMINATING MEMBER:___________________________________Print Name___________________________________ AWS Member No.______________ NOMINATING MEMBER:___________________________________Print Name___________________________________ AWS Member No.______________ NOMINATING MEMBER:___________________________________Print Name___________________________________
CLASSIFIEDS
CAREER OPPORTUNITIES
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Welding Engineering Technology Faculty (9 mos. Full-Time, Tenure Track) Teach undergraduate-level lecture and laboratory course work in a process oriented, “hands-on” A.A.S. and B.S. degree program. Required: Bachelor of Science in Welding Engineering, Welding Engineering Technology, or a closely related field. Two (2) years of welding related experience in a welding application, design, educational, procedure or research environment. Candidates must demonstrate proficiency in GMAW, SMAW, GTAW, OFW, OFC, PAC, SAW, FCAW and RSW. Additional requirements include knowledge of pipe welding and experience in welding graphics, welding fabrication, destructive and nondestructive weldment evaluation, mechanical testing, and computer applications. The successful candidate will have a Master’s degree by the time of appointment or will be required to obtain such a degree within four (4) years of hiring. For a complete posting or to apply, access the electronic applicant system by logging on to http://employment.ferris.edu. Ferris State University is sincerely committed to being a truly diverse institution and actively seeks applications from women, minorities, and other underrepresented groups. An Equal Opportunity/Affirmative Action employer.
3-ton through 120-ton rolls www.joefuller.com
email: [email protected] Phone: (979) 277-8343 Fax: (281) 290-6184 Our products are made in the USA
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Also offering: 9–Year CWI Recertification, RT Film Interpretation, MT/PT/UT Thickness, Welding Procedure Fundamentals, CWS, SCWI, Advanced Inspection Courses
[email protected] (800) 218-9620 (713) 943-8032 WELDING JOURNAL
95
ADVERTISER INDEX ALM Materials Handling . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .10 www.almmh.com . . . . . . . . . . . . . . . . . . . . . . . . . . .(800) 544-5438
Harris Products Group . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .29 www.harrisproductsgroup.com . . . . . . . . . . . . . . .(800) 733-4043
ArcOne . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .41 www.Arc1Weldsafe.com . . . . . . . . . . . . . . . . . . . . . .(800) 223-4685
Hobart Inst. of Welding Technology . . . . . . . . . . . . . . . . . . . . . .60 www.welding.org . . . . . . . . . . . . . . . . . . . . . . . . . . .(800) 332-9448
Arcos Industries, LLC . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .IBC www.arcos.us . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .(800) 233-8460
Image of Welding/WEMCO . . . . . . . . . . . . . . . . . . . . . . . . . . . . .57 www.aws.org/wemco . . . . . . . . . . . . . . . . . . . . . . . .(800) 443-9353
Astaras Welding Accessories, Inc. . . . . . . . . . . . . . . . . . . . . . . .88 www.e3tungsten.com . . . . . . . . . . . . . . . . . . . . . .web contact only
Intercon Enterprises, Inc. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .47 www.intercononline.com . . . . . . . . . . . . . . . . . . . . .(800) 665-6655
Atlas Welding Accessories, Inc. . . . . . . . . . . . . . . . . . . . . . . . . . .26 www.atlaswelding.com . . . . . . . . . . . . . . . . . . . . . .(800) 962-9353
Lincoln Electric Co. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .OBC www.lincolnelectric.com . . . . . . . . . . . . . . . . . . . . .(216) 481-8100
AWS Education Services . . . . . . . . . . . . . . . . . . . . . .59, 64, 87, 91 www.aws.org/education/ . . . . . . . . . . . . . . . . . . . . .(800) 443-9353
Masterweld Products USA . . . . . . . . . . . . . . . . . . . . . . . . . . . . .54 www.masterweld.net . . . . . . . . . . . . . . . . . . . . . . . .(800) 433-3454
AWS Membership Services . . . . . . . . . . . . . . . . . . . . . . . . . .63, 89 www.aws.org/membership/ . . . . . . . . . . . . . . . . . . .(800) 443-9353
Midalloy . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .26 www.midalloy.com . . . . . . . . . . . . . . . . . . . . . . . . . .(800) 776-3300
Bellingham Technical College Welding Rodeo . . . . . . . . . . . . .27 www.btc.cdc.edu . . . . . . . . . . . . . . . . . . . . . . . . . . . .(360) 752-7000
National Bronze & Metals, Inc. . . . . . . . . . . . . . . . . . . . . . . . . .28 www.nbmmetals.com . . . . . . . . . . . . . . . . . . . . . . . .(713) 869-9600
Camfil Air Pollution Control . . . . . . . . . . . . . . . . . . . . . . . . . . . .2 www.camfilapc.com . . . . . . . . . . . . . . . . . . . . . . . . .(800) 479-6801
Netbraze LLC . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .11 www.netbraze.com . . . . . . . . . . . . . . . . . . . . . . . . . .(855) 444-1440
Champion Welding Alloys . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .90 www.championwelding.com . . . . . . . . . . . . . . . . . .(800) 321-9353
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96
MAY 2013
SUPPLEMENT TO THE WELDING JOURNAL, MAY 2013
Sponsored by the American Welding Society and the Welding Research Council
Prediction of -Phase Embrittlement in Type 316FR Weld Metal Precipitation behavior and impact toughness were studied for 316FR weld metal subjected to the high operating temperatures and sodium environment of fast breeder reactors BY E. J. CHUN, H. BABA, K. TERASHIMA, K. NISHIMOTO, AND K. SAIDA
ABSTRACT Through aging treatments at 873 to 1023 K, sigma- ( σ ) phase embrittlement in Type 316FR stainless steel weld metal was predicted at service-exposure temperatures (773–823 K) of a fast breeder reactor (FBR) based on a kinetic approach to σ -phase precipitation. Microstructural examination by scanning and transmission electron microscopies (SEM and TEM) revealed that the dominant precipitated phases were σ and chi (χ), nucleated at δ-ferrite/austenite ( γ) interfaces or in the interior of the δ-ferrite grains, thereby consuming the δ-ferrite during isothermal holds at each aging temperature. The total amount of precipitated phases during isothermal aging sigmoidally increased as a function of the aging time. The kinetics of the appearance of these intermetallic phases could be expressed approximately by a Johnson-Mehl type equation. Based on the determined kinetic equation, the precipitation behavior of intermetallic phases and the degradation of impact toughness at 773 and 823 K could be successfully predicted.
Introduction Recently, society’s most important mission and goal is to secure an effective energy source that will contribute to a reduction in global warming and replacement of the fossil fuels that are gradually running short worldwide. Among various possible energy sources, one that is increasingly coming into the spotlight is nuclear power generation, due to its merits of providing a stable and efficient energy supply system despite its known liabilities of dangerous radioactivity and nuclear waste concerns. Among the diverse types of nuclear power generation, the fast breeder reactor (FBR, using fast neutrons that breed Pu-239 from U-238) is well known as the most advanced thanks to its superior fuel economy. Thus, much research has been performed in Japan on the metallurgical behavior, as well as the structural properties, of materials for the Japanese prototype fast breeder reactor E. J. CHUN, H. BABA, K. TERASHIMA, and K. SAIDA are with Division of Materials and Manufacturing Science, Graduate School of En gine erin g, Osak a Unive rsity, Osak a, Japa n. K. NISHIMOTO is with Department of the Ap plication of Nuclear Technology, Fukui University of Technology, Fukui, Japan.
Monju from the 1970s to the present day (Refs. 1–5). For future commercial FBR plants, further research on welding materials will need to be performed. In addition, as present nuclear plants have been in operation for a long time, repairs of their aging parts, which mainly involve welding, are becoming necessary. Prior to discussing repair welding processes for the FBR, it is important to understand the typical differences between a FBR and other general types of reactors (i.e., light water-cooled reactor). First, a FBR cannot use water as a moderator, rather molten sodium metal is employed as a coolant, because of its high thermal conductivity. Second, a FBR has a higher operating temperature of 773–823 K compared to other reactors, due to its high degree of heat generation (Ref. 4). Therefore, before the commer-
KEYWORDS Embrittlement Type 316FR Weld Metal Sigma Phase Kinetics
cialization of future FBRs in Japan can become a reality, discovering the effects of the sodium environment and the relatively high operating temperature on the repair weldability of components made of austenitic stainless steel (Types 304, 316, and 321), the alloy usually chosen due to its superior corrosion resistance, ductility, strength, formability, and weldability (Refs. 6, 7), are very significant issues. Furthermore, an advanced 316FR stainless steel structural material, which has improved creep fatigue behavior over other austenitic stainless steels, and which possesses a higher phase stability during high temperature at longer exposure times by using the concept of solid-solution hardening with low-carbon and medium nitrogen as compared to conventional austenitic stainless steel, is likely to be used for the next generation of commercial FBRs in Japan. However, to the best of the authors’ knowledge, research on these issues has not been reported to date. A particular result of the high-temperature operation of a FBR is the need for weld repairs to the various main com ponents in the nuclear plant. Such problems arise because welds have poorer quality than the base metals due to solute segregation or microstructural inhomogeneity (to avoid hot cracking, the weld metal for austenitic stainless steel is often intentionally rendered inhomogeneous by introducing some amount of δ−ferrite as a result of the rapid solidification rate during the welding process). These issues promote the transformation of intermetallic phases, which affect various mechanical and chemical properties (Ref. 8). Therefore, in anticipation of the need for welding repairs to FBRs, the study and prediction of the aging behavior of the weld metal is an absolutely necessary prerequisite. Consequently, as a first research step toward developing a process for weld repairs in the newly developed 316FR stainWELDING JOURNAL 133-s
H C R A E S E R G N I D L E W
W Fig. 1 — Schematic illustration of the machining and detailed dimensions E of specimens for the Charpy impact test. L D less steel needed for the future generation AF solidification the objective of the present study mode, solidifying as I N ofis toFBRs, clarify the prediction of -phase emprimary austenite ( ), G brittlement behavior. This is achieved which is generally rethrough study of the precipitation kinetics garded as having a R of the -phase in the weld metal at the pronounced high-tem2 — SEM micrographs o f a cross section of b ead-on-plate one-pass E general service exposure temperature of a perature cracking Fig. weld metal. In particular, the weld metal of typsusceptibility. S FBR. austenitic stainless steels can contain E ical different final microstructure characterisand A tics that are determined by solidification Materials plate material for the Charpy impact tests. Experimental Procedures The chemical compositions of these mateR and subsequent phase transformation, typrials are given in Table 1. through the AF and FA solidification C ically Materials (Refs. 8, 9). Different solidification H modes Aging Treatment of Weld Metal modes strongly affect various properties σ
γ
σ
such as corrosion resistance, high-temperature cracking behavior, and microstructural properties, etc. (Refs. 8–12). In this study, we first examined weld metal in the
The base metal used in this study for aging treatments and Charpy impact tests was Type 316FR austenitic stainless steel. The 316L types were only employed by the
Prior to performing aging treatments, bead-on-plate welds were prepared using a gas tungsten arc welding (GTAW)
Table 1 — Chemical Compositions of Materials Used (mass-%)
Type 316FR Type 316L
C
S
P
Cr
Ni
Mo
Si
Mn
N
Al
O
Fe
0.0085
0.0009
0.023
17.56
12.02
2.15
0.44
0.79
0.088
0.011
0.007
Bal.
0.0150
0.0040
0.031
17.28
12.11
2.04
0.70
0.92
0.024
—
—
Bal.
— 1000 — —
1127 — — —
Table 2 — Aging Conditions Used in the Present Study
Temperature (K) 873 923 973 1023
Aging Time (h) 0.5 0.5 0.5 0.5
134-s MAY 2013, VOL. 92
1 1 1 1
5 5 5 5
10 10 10 10
50 50 50 50
100 100 100 100
394 — — —
— 500 500 —
526 — — —
— 1532 — —
process. The dimensions were 120 × 40 × 3 mm. Detailed GTAW conditions were as follows: arc current, 110 A; arc voltage, 14 V; and welding speed, 1.67 mm/s. We specimens were aged by heating as indicated in Table 2, and after the aging treatment, they were quenched with water. Microstructure Analysis
The content of δ− ferrite in the as welded and aged specimens was measured using the magnetic induction method (Feritscope®) in the central area of the bead surface. To clarify microstructural changes caused by the aging treatments, specimens were observed by scanning electron microscopy (SEM) equipped with electron backscatter diffraction (EBSD) under an acceleration voltage of 20 kV after electrolytic etching with a 10% aqueous solution of KOH using an applied voltage of 100 mV. The specimens were also observed by transmission microscopy (TEM) under an acceleration voltage of 200 kV after jet polishing with a solution of perchloric acid (5%) and acetic acid (95%), using an applied voltage of 50 V. Charpy Impact Test
Details of the machining and the dimensions of welding specimen are indicated in Fig. 1. Smaller Charpy impact test specimens of size 55 × 10 × 3 mm were machined from welded plate as also indicated in Fig. 1. The aging treatment for Charpy impact test specimens was heating at 1023 K for 0, 0.5, 1, and 10 h to differentiate δ−ferrite decomposition behavior, while the tests themselves were conducted on the as-welded and aged samples at room temperature according to JIS Z 2242, Method for Charpy pendulum impact test of metallic materials . Four specimens were tested for each aging time and their average taken as the absorbed impact energy value. After the impact test, the fractured surfaces of the specimens were observed using SEM.
Change in Microstructure with Aging Treatment Figure 2 shows typical microstructures for Type 316FR weld metal after the GTAW process. The cell morphology associated with the solidification behavior is clearly visible. All the δ-ferrite was located at the cell boundaries or triple points of austenite ( γ) with elongated or globular shape. These δ-ferrite distributions were generally identified as being diagnostic of the AF solidification mode, and the average volume fraction of δ-ferrite was about 3 % (FN < 3), measured by a Feritscope. Figure 3 shows SEM micrographs of
Fig. 3 — SEM micrographs of the weld metal after aging at 873 K for 10 and 100 h, and 1023 K for 1 h.
H C R A E S E R Fig. 4 — TEM micrographs and diffraction patterns of intermetallic phases ( σ and χ phases). G N I precipitation behavior during aging at 873 tified as the sigma phase ( σ : FeCr) and the D and 1023 K for various holding times. The chi phase ( χ: Fe18Cr6Mo5) through bright L precipitates appeared in the interior of the and dark images, selected area diffraction E δ-ferrite, and increased in number with an pattern, and its key diagram analysis. It increase in aging time. The precipitates follows that the predominant precipitates W were mainly classified as belonging to two types, one nucleated at the δ-ferrite/ γ interface and one within the δ-ferrite. These precipitates can be easily identified on the basis of backscattered electrons (BSE) contrast in SEM. As reported elsewhere, these precipitation behaviors may totally consume the δ-ferrite. Figure 4 presents TEM micrographs that reveal two types of precipitates iden-
in 316FR stainless steel weld metal during long-term aging were σ and χ phases. Based on these results, EBSD analysis was performed. Figure 5 shows representative BSE micrographs of SEM and EBSD micrographs. Therefore, one type nucleated at the δ-ferrite/ γ interface was σ -phase and another within the δ-ferrite was χ-phase in Type 316FR weld metal during aging treatment.
Table 3 — Kinetic Parameters Determined for the Prediction of Aging Behavior
Determined Prediction Parameters Temperature (K) k (/s)
873 923 973 1023
1.2 × 10–6 2.9 × 10–6 3.0 × 10–5 1.8 × 10–4
n
k0 (/s)
Q (kJ/mol)
0.320
2.1 × 109
258
WELDING JOURNAL 135-s
Fig. 5 — Backscattered electron micrographs of SEM and EBSD micrographs of the weld metal after aging at 873 K for 100 h.
Fig. 6 — Fractional change of decomposed δ-ferrite with aging time at various aging temperatures.
where f i is the initial amount of the δ-ferrite prior to aging and f f is final amount of the δ-ferrite after aging. Based on this relationship, Fig. 6 shows the change in the decomposed δ-ferrite fraction as a function of the aging temperatures and holding times. The fraction of the intermetallic phases increased sigmoidally with an increase in the aging time at any aging temperature, and approached the saturation point for long-term aging at 973 and 1023 K.
W E L D I N G R E S E A R C Fig. 7 — Applicability of kinetic approaches for the decomposition behavior of δ-ferrite. H Prediction of Intermetallic Phase Precipitation In order to predict the precipitation of intermetallic phases in Type 316FR weld metal during the practical operation of a FBR, we have adopted a kinetic approach to the isothermal aging treatments. Effect of Aging Conditions on Intermetallic Phases Precipitation
Fig. 3, σ - and χ -phases were dominantly precipitated in the δ-ferrite grains, thereby decomposing the δ-ferrite. They are also representative intermetallics, which is associated with degradation of impact toughness. Therefore, the amount of intermetallic phases both σ and χ precipitated could be approximated by the decomposed fraction of δ-ferrite ( y) expressed as follows: y
As mentioned above with reference to
f f =
−
f i
f i
Table 4 — Activation Energy for Diffusion of Alloying Elements in δ-Ferrite
Activation Energy (kJ/mol) Diffusion of Cr in δ−ferrite Diffusion of Ni in δ−ferrite Diffusion of Mo in δ−ferrite Diffusion of Fe in δ−ferrite
136-s MAY 2013, VOL. 92
Kinetic Equation for the δ-Ferrite Decomposition
267 262 283 296
(1)
In order to determine the kinetics of decomposition in δ-ferrite, we employed three approaches commonly used to describe the phase transformation kinetics in various nucleation and growth systems (Refs. 15–22): a parabolic law (the diffusion-controlled growth theory proposed originally by Zener), and the JohnsonMehl and Austin-Rickett equations. In the first case, the parabolic law is expressed by
y
=
k t
(2)
where y is the decomposed fraction of δferrite indicated in Equation 1, k is the parabolic rate constant, and t is the exposure time (aging time in the present study). On the other hand, the kinetic equation of phase transformations in metals can be also generally expressed as dy dt
=
k nt n
1
−
(1
−
y)
m
(3)
where y is the decomposed fraction of δ ferrite indicated in Equation 1, k is the
Fig. 8 — Plot to determine the activation energy for Fig. 9 — Prediction of the decomposition behavior Fig. 10 — C func impact tou ghness as a func— C hange of impact tou δ -ferrite in the weld metal at practical operating decomposition of δ-ferrite at various aging temper- of δ rrite (aging (aging decomposed δ-fe rrite tio n of the amount of amount of decomposed temperatures of a FBR. atures. ). at 1023 K
temperature-dependent rate constant, t is the time (ageing time in the present study), n is the time exponent parameter calculated by regression analysis depending on the nucleation mechanism and the growth processes, and m is the impingement exponent. Equation 3 with m = 1 and m = 2 corresponds to the JohnsonMehl equation (Equation 4) and the Austin-Ricket Aust in-Rickettt equation (Equation (Equation 5): n
y = 1 − exp − ( kt ) y 1
−
y
=
(4)
( k t t )n
Johnson-Mehl plot in Fig. 7. Moreover, it Johnson-Mehl is well known that the dependence of the precipitation rate k in Equation 4 can be generally expressed by an Arrhenius equation as follows:
−Q RT
k = k 0 exp
(6) where k0 is the frequency factor; Q is the activation energy; T is the temperature; and R is the gas constant. Equation 6 can be transformed into the following to describe an Arrhenius plot: ln k = −
(5)
The applicability of these kinetic equations can be confirmed by parabolic, John y)) = n log t+ n son-Meh son -Mehll (log ln ( 1/1– y log k) and Austin-Rickett (log ( y /(1– y) = n log t+ n log k) plots. Figure 7 shows the results of applying the three kinds of kinetic approaches. There was a good linear relationship between the aging time and the fraction of decomposed δ-ferrite in the Type 316FR weld metal in the Johnson-Mehl plot regardless of aging temperature, while the other plots failed to adequately describe the δ-ferrite decomposition behavior. In particular, the Austin-Rickett plot deviated from a linear relationship in the final stages of precipitation, and the parabolic plot showed also totally nonlinear behavior at every aging temperature. In other words, δ-ferrite decomposition behavior during aging of Type 316FR weld metal was best described by the Johnson-Mehl kinetic equation. To predict the decomposition behavior during long-term service exposure of a FBR, the remaining constants in the Johnson-Mehl equation ( n and k) need to be determined. These values were found from a simple regression analysis of the
Q 1 R T
+ 1nk 0
(7) Figure 8 shows the Arrhenius plot, showing a simple linear relationship between the reciprocal temperature and the logarithm of k, allowing the determination of the activation energy for the decomposition of δ-ferrite and the k0 constant. All the parameters determined in this way are shown shown in Table Table 3. The fitting fitting constant n was independent independent of the aging aging temperature, but the precipitation rate constant k increased as a function of aging temperature. Literature report (Ref. 23) of the acti vation energy energy for diffusion diffusion of the the main alalloying elements present within the δ-ferrite is shown in Table 4. Their similar activation energies suggest that the decomposition of δ-ferrite in this study would be strongly strongly influenced influenced by the the diffusion of these alloying elements. Prediction of δ-Ferrite Decomposition at Service-Exposure Temperatures of a FBR
Using the determined kinetics of the Johnson-Mehl equation, the decomposition behavior of δ-ferrite during in-ser vice exposure of FBR at practical operation temperatures (773 and 823 K) was predicted. Figure 9 shows the predic-
tion results, showing a sigmoidal relationship between operating time and the decomposed fraction of δ-ferrite. There was a large difference in the decomposition behavior between operating temperatures of 773 and 823 K. Specifically, it took 15 months to reach 50% δ-ferrite decomposition at 773 K, but it took only 1 month at an operating temperature of 823 K to arrive at the same point. In other words, a 50 K difference in the operating temperature caused about a 15 times faster decomposition rate. The decomposed fraction of δ ferrite at other points on the aging time curve is also listed in Table 5. According to these predictions, about 90% of the δ-ferrite will have decomposed after only 5 years of operation of a FBR at 823 K, while about 66% of the δ-ferrite will have decomposed at an operating temperature of 773 K. Further Furthermore, more, after 50 years of FBR operation, o peration, in Type Type 316FR weld metal, δ -ferrite will be completely decomposed at 823 K. Consequently, the operating temperature of the FBR should be closely considered from the viewpoint of δ δ -ferrite decomposition behavior, because it could seriously affect not only various mechanical and chemical properties during service, but also the weldability of any required repairs. Prediction of Impact Toughness Behavior in Weld Metal
In order to predict the embrittlement behavior resulting from the decomposition of δ-ferrite in Type 316FR weld metal during practical operation of a FBR, Charpy impact tests were performed for specimens with different different fractions fractions of decomp decomposed osed δferrite. To predict this behavior, one assumption that impact toughness of weld metal is only governed by fractional change of δ-ferrite decomposition was employed in this study.
WELDING JOURNAL 137-s
H C R A E S E R G N I D L E W
Fig. 11 — SEM fractographs after Charpy impact tests as a function of decomposed fraction of δ-ferrite Fig. 12 — Predicted results of the absorbed impact (aging at 1023 K). energy during long-term operation of a FBR.
W Effect of δ-Ferrite Decomposition on E Impact Toughness L D Figure 10 shows the variation of imI absorbed energy as a function of N pact aging time (i.e., decomposed fraction of G δ-ferrite). The impact energy of the weld gradually decreased with increasR metal ing isothermal holding time, namely, the E absorbed energy decreased with an inS crease in the fraction of decomposed δferrite. Particularly, the energy value of E weld metal aged for 10 h was about 22 J, A a drastic by a factor of two comR pared withdecline an as-welded specimen. This C is the well-known σ -phase -phase embrittlement in stainless steel weld metal (Refs. H found 6, 7, 19, 24). Figure 11 compares SEM fractographs of as-welded and aged weld metal after the impact test to characterize the differences in failure mechanism. In the as-welded specimen, all the failure essentially occurred in the shallow dimpled fracture mode. However, in the aged specimens (decomposed fraction of δ-ferrite: 0.35, 0.60, and 0.90), some brittle fracture area was also detected detected in the surface, surface, although the majority of the surface had suffered
shallow dimpled fracture surface. The fraction of brittle fracture also increased with an increase in the decomposed fraction of δ-ferrite. In particular, as the isothermal holding time increased, the fractographs of the brittle fracture areas revealed that the fracture surface was largely composed of a dendrite morphology. These fractographs show that the brittle fracture was initiated in a dendrite region by the precipitation of intermetallic phases ( σ - and χ-phases) in δ-ferrite. Thus, it was clear that the presence of intermetallic phases significantly decreased the impact toughness of the weld metal. Predicted Impact Test Results at the Service-Exposure Service-Expos ure Temperature of a FBR
By substituting the results of the Charpy impact tests shown in Fig. 10 into the predictions of precipitation behavior shown in Fig. 9, we were able to predict the impact toughness during long-term service exposure in a FBR. Figure 12 shows the predicted results of the absorbed impact energy. As the result of the precipitation behavior shown in Fig. 8, the absorbed impact energy also de-
Table 5 — Pred Predicted icted Results of the Decomposed Fraction Fra ction of of δ-Ferrite at Some Major Time Pointss in Fig. 9 Point
Decomposed Fract Frac tion of δ−Ferrite Operating time (years) 5 10 30 50
138-s MAY 2013, VOL. 92
773 K 0. 6 6 0.74 0.85 0.89
8 23 K 0.90 0.95 0.98 0.99
creased sigmoidally with an increase in the operating time. It takes approximately 3 months to reach an absorbed impact energy of 32 J at 823 K, while approximately 32 months will be needed to reach the same absorbed energy at 773 K. Therefore, it was clear that impact toughness behavior is also affected by differences in the operating temperature.
Conclusions In this study, the precipitation behavior and changes in impact toughness during actual operation of FBR was predicted for Type 316FR stainless steel weld metal, based on the kinetics of δ -ferrite decomposition. The main conclusions of this work can be summarized summarized as follows: 1) Type 316FR stainless steel weld metal contained approximately 3% (FN < 3) of δ-ferrite formed in the AF solidification mode. Intermetallic σ - an and d χ- ph phase asess were precipitated inside δ-ferrite during aging treatments at 823, 873, 923, and 1023 K, consuming the δ-ferrite (i.e., fraction of decomposed δ-ferrite = fraction of precipitated intermetallic phases). As the aging temperature and isothermal hold time increased, the amount of intermetallic phases increased sigmoidally. 2) The decomposition of δ-ferrite was examined by three types of kinetic approaches — the parabolic law, Austin-Rickett, and Johnson-Mehl equations. Among these kinetic approaches, the decomposition behavior was best described by the Johnson-Mehl type equation. The kinetic parameters in the Johnson-Mehl type equation were determined to be n = 0.320, Q = 258 kJ/mol, and k0 = 2.1 × 109 /s, rega regardl rdless ess of the aging temperature.
3) The δ-ferrite decomposition behavior at practical operating temperature (773 and 823 K) found during FBR service was predicted using the above determined parameters. The decomposition rate of δ-ferrite at an 823 K service temperature was 15 times faster than that at 773 K. 4) As the isothermal holding time increased, the impact toughness decreased during the aging treatment at 1023 K. In the fractographs obtained after impact tests, the fraction of brittle fracture region increased with an increase in decomposed δ-ferrite. 5) By combining the predictions of δ ferrite decomposition behavior with the experimental results of the Charpy impact test, we were able to predict changes in impact toughness during in-service of a FBR. We predict that it will take approximately 3 months to reach an absorbed impact energy of 32 J at 823 K, while approximately 32 months will be needed to reach a the same absorbed energy at 773 K. In other words, the rate of decrease of the absorbed impact energy at an 823 K operating temperature was 10 times faster than that at 773 K. Acknowledgment
The present study includes the result of “Core R&D program for commercialization of the fast breeder reactor by utilizing Monju” entrusted to University of Fukui by Ministry of Education, Culture, Sports, Science and Technology of Japan (MEXT). References 1. Furukawa, T., Kato, S., and Yoshida, E. 2009. Compatibility of FBR materials with sodium. Journal of Nuclear Materials Materials 392(2): 249 to 254. 2. Iida, K., Asada, Y., Okabayashi, K., and Nagata, T. 1987. Construction codes developed for prototype FBR Monju. Nuclear Engineering and Design 98 (3): 283 to 288. 3. Iida, K., Asada, Y., Okabayashi, K., and
Nagata, T. 1987. Simplified analysis and design for elevated temperature components of Monju. Nuclear Engineering and Desig n 98(3): 305 to 317. 4. Nakazawa, T., Kimura, H., Kimura, K., and Kaguchi, H. 2003. Advanced type stainless steel 316FR for fast breeder reactor structures. Journal of Material s Processing Technolog y 143–144(20): 905 to 909. 5. Nogami, S., Hasegawa, A., Tanno, T., Imasaki, K., and Abe, K. 2011. High-temperature helium embrittlement of 316FR steel. Journa Journall of Nuclear Scienc Sciencee and and Technolo echnology gy 48(1): 130 to 134. 6. Gill, T. P. S., Vijayalakshmi, M., Rodriguez, P., and Padmanabhan, K. A. 1989. On microstructure-property correlation of thermally aged Type 316L stainless steel weld metal. Metallurg ical and Material s Transactions A 20(6): 1116 to 1124. 7. Ibrahim, O. H., Ibrahim, I. S., and Khalifa, T. A. F. 2010. Effect of aging on the toughness of austenitic and duplex stainless steel weldments. Journal of Material s Science and Technology 26(9): 810 to 816. 8. Inoue, H., Koseki, T., Okita, S., and Fuji, M. 1997. Solidification and transformation behavior of austenitic stainless steel weld metals solidified as primary austenite: Study of solidification and subsequent transformation of CrNi stainless steel weld metals — 1st Report. Weldingg International 11(11): 876 to 887. Weldin 9. Inoue, H., Koseki, T., Okita, S., and Fuji, M. 1997. Solidification and transformation behavior of austenitic stainless steel weld metals solidified as primary ferrite: Study of solidification and subsequent transformation of Cr-Ni stainless steel weld metals — 2nd Report. Welding International 11(12): 937 to 949. 10. Inoue, H., Koseki, T., Okita, S., and Fuji, M. 1998. Epitaxial growth and phase formation of austenitic stainless steel weld metals near fusion boundaries: Study of solidification and transformation of Cr-Ni stainless steel weld metals — 3rd Report. Welding International 12(3): 195 to 206. 11. Inoue, H., Koseki, T., Okita, S., and Fuji, M. 1998. Solidification and transformation behavior of Cr-Ni stainless steel weld metals with ferritic single phase solidification mode: Study of solidification and transformation of Cr-Ni stainless steel weld metals — 4th report. Welding International 12(3): 282 to 296. 12. Kou, S. 2003. Welding Metallurgy, second edition. P. 216-232, A John Wiley & Sons, Inc. 13. Hsieh, C. C., and Wu, W. W. 2012. Overview
of intermetallic sigma ( σ ) phase precipitation in stainless steels. ISRN Metallurgy 2012: Article ID 732471, 16 pages. 14. Escriba, D. M., Materna-Morris, E., Plaut, R. L., and Padilha, A. F. 2009. Chi-phase precipitation in a duplex stainless steel. Materi Materials als CharChar acterization acteriz ation 60(11): 1214-s to 1219-s. 15. Wang, Z., Mao, X., Yang, Z., Sun, X., Yong, Q., Li, Z., and Weng, Y. 2011. Straininduced precipitation in a Ti micro-alloyed HSLA steel. Materials Science and Engineering A 529(25): 459 to 467. 16. Starink, M. J. 1997. Kinetic equations for diffusion-controlled precipitation reactions. Journal of Material s Science 32(15): 4061 to 4070. 17. Badji, R., Bouabdallah, M., Bacroix, B., Kahloun, C., Bettahar, K., and Kherrouba, N. 2008. Effect of solution treatment temperature on the precipitation kinetic of σ -phase -phase in 2205 duplex stainless steel welds. Material s Scie nce and Engineering A A 496(25): 447 to 454. 454. 18. Johnson, W. A., and Mehl, R. F. 1939. Reaction kinetics in processes of nucleation and Growth. Transactions AIME 135: 416 to 458. 19. Sasikala, G., Ray, S. K., and Mannan, S. L. 2003. Kinetics of transformation of delta ferrite during creep in a Type 316(N) stainless steel weld metal. metal. Materials Science Science and Engineering A 359 (25): 86 to 90. 20. Sello, M. P., and Stumpf, W. E. 2011. Laves phase precipitation and its transformation kinetics in the ferritic stainless steel type AISI 441. Materials Science and Engineering A 528(25): 1840 to 1847. 21. Lee, E. S., and Kim, Y. G. 1990. A transformation kinetic model and its application to Cu-Zn-Al shape memory alloys — 1. isothermal Metallurgica et Materialia Materialia 38(9): conditions. Acta Metallurgica 1669 to 1676. 22. Sutou, Y., Koeda, N., Omori, T., Kainuma, R., and Ishida, K. 2009. Effects of aging on bainitic and thermally induced martensitic transformations in ductile Cu-Al-Mn-based shape memory alloys. Acta Material ia 57(19): 5748 to 5758. 23. Magnabosco, R. 2009. Kinetics of sigma phase formation in a duplex stainless steel. Materials Research 12(3): 321 to 327. 24. Lee, D. J., Byun, J. C., Sung, J. H., and Lee, H. W. 2009. The dependence of crack properties on the Cr/Ni equivalent ratio in AISI 304L austenitic stainless steel weld metals. Materials Science and Engineering A 513–514(15): 154 to 159.
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Microstructural Evolution and Mechanical Properties of Simulated Heat-Affected Zones in an Iron-Copper Based Multicomponent Steel A combination of dilatometry, HAZ simulations, and mechanical testing were used to determine the mechanical properties that develop in the HAZ of NUCu-140 BY J. D. FARREN, A. H. HUNTER, J. N. D UPONT, C. V. ROBINO, E. KOZESCHNIK, AND D. N. SEIDMAN
ABSTRACT
W E L D I N G R E S E A R C H
NUCu-140 is a recently developed steel that relies on nano-scale Cu-rich precipitates to achieve yield strength levels in excess of 825 MPa (120 ksi). In order for NUCu-140 to be utilized as a structural material, a comprehensive welding strategy must be developed. Since NUCu-140 is a precipitation-strengthened material, this strategy must include a detailed understanding of the precipitate evolution that occurs in the heat-affected zone (HAZ) as a result of welding thermal cycles. A combination of dilatometry, HAZ simulations, and mechanical testing are presented to determine the mechanical properties that develop in the HAZ of NUCu-140. MatCalc kinetic simulations and Russell-Brown strengthening calculations were conducted to model the observed precipitate and mechanical property trends. The microhardness and tensile testing results reveal that local softening is expec ted in the HAZ of NUCu140 welds. MatCalc simulations show that a combination of partial dissolution, full dissolution, and re-precipitation of the Cu-rich precipitates is expected to occur in the various HAZ regions. The predicted precipitate parameters are used as input to the Russell-Brown strengthening model to estimate the changes in strength expected due to changes in precipitate features. The measured and predicted strength levels exhibit very good quantitative agreement for the low-heat-input simulations and reasonable qualitative agreement for the high-heat-input weld simulations.
Introduction Copper precipitation-strengthened materials such as high-strength, low-alloy (HSLA) 80 and 100 have been used extensively in naval and structural applications due to their excellent combination of strength and toughness. As a result of the ever increasing need to minimize cost, it is desirable to develop a HSLA variant that can achieve even higher yield strength lev J. D. FARREN is with Naval Surfac e Warfare Center – Carderock Division, West Bethesda, Md. A. H. HUNTER and D. N. SEIDMAN are with Dept . of Mate rial s Scie nce and Engi neering, Nort hwestern Unive rsity, Evan ston, Ill. J. N. DU PONT ([email protected]) is with Dept. of Materials Science and Engineering, Lehigh Uni versity, Bethlehem, Pa. C. V. ROBINO is with San dia National Laboratories, Albuquerque, N.Mex. E. KOZESCHNIK is with Dept. of Materials Sci ence and Technology, Vienna University of Tech nology, Vienna, Austria.
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els [≥ 825 MPa (120 ksi)], while maintaining suitable toughness. Recent research conducted at Northwestern University has produced a candidate structural material that achieves yield strength levels in excess of 825 MPa while retaining toughness levels that would exceed the requirements for most naval and structural applications (Refs. 1–9). NUCu-140 is a copper precipitation-strengthened steel that is composed of a nominally ferritic microstructure with nano-scale Cu-rich precipitates that strengthen the material, and NbC precipitates that limit the austenite grain growth. The use of NUCu-140 can offer
KEYWORDS High-Strength Steels Fracture Weld Process Simulation
significant cost savings as a result of 1) minimization of expensive alloying elements; 2) simple production using inexpensive processing techniques; and 3) construction of structurally sound designs using less material due to higher yield strength. It has been estimated that the utilization of NUCu-140 can produce a fabrication cost savings of 20–35% (Ref. 10). In order for NUCu-140 to be utilized as a structural material, a comprehensive welding strategy must be developed. Since NUCu-140 is a precipitation-strengthened material, this strategy must include a detailed understanding of the precipitate evolution that occurs in the heat-affected zone (HAZ). There are four distinct regions that typically develop in the HAZ of steel welds: 1) the subcritical HAZ region; 2) the intercritical HAZ region; 3) the fine-grained austenite HAZ region; and 4) the coarse-grained austenite HAZ region (Ref. 11). These regions are defined by the peak temperatures that they experience relative to the austenization temperatures of the material, Ac1 and Ac3. The subcritical HAZ experiences a peak temperature during the weld thermal cycle that does not exceed the austenite start temperature (Ac1). Therefore, the subcritical region does not undergo any transformation to austenite. The intercritical region experiences a peak temperature between the Ac1 and Ac3 temperatures that results in partial transformation to austenite during the weld thermal cycle. The fine-grained HAZ region experiences a peak temperature that exceeds the Ac3 temperature and therefore causes full transformation to austenite during welding. In this region, the thermal cycle only minimally exceeds the Ac3 temperature, which prevents significant austenite grain growth. The final HAZ region is the coarse-grained HAZ. In this region, the Ac3 temperature is significantly exceeded, which leads to austen-
Fig. 2 — Dilatometry results for NUCu-140 heated at 1°, 10°, 100°, and 1000°C/s.
Fig. 1 — LEAP tomography data collected from NUCu-140 GMAW showing the evolution of the average precipitate radius (), number den sity, and volum e fraction ( φ ) across the base metal, HAZ, and FZ of the weld.
ite grain growth. The intercritical, finegrained, and coarse-grained HAZ regions all experience transformation to austenite during the weld thermal cycle. Therefore, both the microstructural and precipitate evolution needs to be investigated to understand the mechanical properties in these regions. A preliminary investigation of the microstructural evolution and mechanical properties in NUCu-140 gas metal arc welds (GMAW) and gas tungsten arc welds (GTAW) was recently conducted (Ref. 12). Microhardness traces revealed that a locally softened HAZ region formed as a result of the fusion welding process. Average precipitate radius (), number density (Nv), and volume fraction measurements ( φ) conducted using local electrode atom probe (LEAP) tomography confirmed that the observed decrease in microhardness occurred as a result of the precipitate evolution that occurs in the HAZ. Figure 1 shows a summary of the results. The base metal region shows the initial precipitate parameters that develop as a result of the solution and aging thermal treatment. The region labeled HAZ 1 experienced a peak temperature of ~ 675°C and exhibits a reduction
in and φ while showing a concomitant increase in Nv. This results from partial dissolution of the precipitates on heating, followed by re-precipitation of new and smaller Cu-rich precipitates during cooling. The region labeled HAZ 2 experienced a peak temperature of ~910°C and exhibits a further decrease of the and φ with an even greater increase in Nv. It was determined that full precipitate dissolution occurs in HAZ 2 on heating, followed by re-precipitation on cooling. The fusion zone also undergoes full dissolution of the Cu-rich precipitates on heating but exhibits only minimal re-precipitation during the cooling portion of the weld cycle. Therefore, the fusion zone exhibits the lowest , Nv, and φ of any weld region. The overall trends in , Nv, φ are consistent with the observed local softening that occurs in the HAZ as a result of the fusion welding process. The current research focuses on a more detailed investigation of the mechanical properties of each of the four critical regions of the HAZ using simulated HAZ samples.
Experimental Procedure The chemical composition of the NUCu-140 steel investigated in this study was measured using inductively coupled plasma-optical emission spectroscopy (ICP-OES) and the results are shown in Table 1. The composition of NUCu-140 is similar to HSLA-100 Comp II, with slightly increased C and Al levels and slightly decreased Cr and Mo levels. The Al content is relatively high compared to traditional structural steels, but Al has been shown to segregate to the interface of the Cu-rich precipitates in NUCu-140 and is believed to limit the coarsening kinetics during aging (Ref. 3). The NUCu-
Table 1 — Composition of Copper Precipitation-Strengthened NUCu-140 Steel (all values in wt-%)
Element
NUCu-140
Al C Cu Fe Mn Nb Ni P S Si
0.65 0.04 1.35 Bal. 0.47 0.07 2.75 0.009 0.002 0.47
140 was vacuum melted, cast into ingots, and homogenized at 1150°C for 3 h. The ingots were hot rolled at approximately 950°C to a final plate thickness of 25.4 mm and air cooled. The plates were solutionized at 900°C for 1 h, water quenched to room temperature, aged at 550°C for 2 h, and air cooled to room temperature. Critical transformation temperatures at various heating and cooling rates were determined by dilatometric methods using a Gleeble 3500 thermomechanical simulator equipped with an Anritsu SLB laser dilatometer. Diametral dilatometry was conducted on 6.35-mm-diameter solid samples oriented such that the diameter measurements corresponded with the through-thickness direction in the original plate. Samples were taken from approximately the ¼ thickness position in the plate. A freespan of 25 mm and a lowforce jaw carrier were used to minimize constraint during the determinations. The Ac1 and Ac3 determinations generally followed ASTM A1033, although some deviations from recommended heating rates and conditioning temperatures were used. WELDING JOURNAL 141-s
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A
Table 2 — Summary of the P recipitate Evolution Predicted Using MatCalc Kinetic Simulation Software
Sample
B
Fig. 3 — Summa ry of the therm al cycle s pre dicted using SOAR for the foll owing sam ples: A — Low heat input (1.5 J/m); B — high heat input (3.75 J/m).
W E L D I N G R E S E A R C Fig. 4 — LOM micrograph showing the NUCuH 140 base metal microstructure. In order to be more consistent with the NUCu140 aging temperature, the samples were preconditioned at 450°C rather than 600°C. Preliminary experiments indicated that Ac1 in this steel is below 700°C, so the heating rate change for the Ac3 determinations was conducted at 590°C rather than the 700°C recommended by ASTM A1033. This information was then used to formulate an experimental matrix of simulated HAZ samples that would represent each of the four critical regions of the HAZ. The HAZ simulations were conducted using a Gleeble 3500 thermomechanical simulator, and the heating and cooling rates were controlled using the QuikSim software package supplied with the Gleeble 3500. Thermal cycles associated with a high (3.75 kJ/mm) and low (1.5 kJ/mm) heat input were utilized to represent the range of arc welding conditions expected during the joining of NUCu-140. 142-s
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Base Metal (BM) LH675 LH800 LH900 LH1350 HH675 HH800 HH900 HH1350
Peak Temperature (°C)
MatCalc Radius (nm)
MatCalc Number Density (m–3)
MatCalc Phase Fraction
150 675 800 900 1350 675 800 900 1350
4.25 1.11 0.56 0.54 0.46 0.96 0.60 0.60 0.60
2.53 × 1022 1.10 × 1023 5.86 × 1024 5.95 × 1024 7.51 × 1024 1.24 × 1023 5.19 × 1024 5.17 × 1024 5.19 × 1024
0.0099 0.0064 0.0042 0.0041 0.0030 0.0068 0.0051 0.0051 0.0051
The simulated HAZ samples were taken in the T-L orientation and were 11 × 11 × 60 mm. The samples were outfitted with multiple thermocouples to determine the width of the uniformly heated region within the freespan (10 mm). Specimens were prepared for light optical microscopy (LOM) using standard metallographic techniques and etched using a 3% Nital solution. Grain size measurements were conducted according to ASTM Standard E112-96(2004)ε 2, and five fields were measured per sample. Microhardness measurements were conducted using a Vickers diamond indenter, a 1-kg load, and a 15-s dwell time. Charpy impact testing was conducted at –40°C according to ASTM E23 on 10- × 10- × 50-mm specimens. Tensile testing was performed according to ASTM E8 on subsized samples with a diameter of 6.35 mm and a 25.4-mm gauge length. Charpy impact testing was performed on two specimens per condition while tensile testing was conducted on a single specimen per condition due to material limitations. Post-fracture analysis was conducted on all of the Charpy impact and tensile specimens to ensure that crack propagation and subsequent failure occurred within the uniformly heated region of the simulated HAZ sample. If failure was not contained entirely within the uniformly heated region, the results of the sample were discarded and a replacement specimen was prepared and tested. Results and Discussion
In order to simulate the four critical regions of the HAZ, it is first necessary to identify the critical transformation temperatures, Ac1 and Ac3, for the NUCu140 substrate material. Dilatometry heating rates ranging from 1° to 1000°C/s were investigated and the results are shown in Fig. 2. The dilatometry curves show that the Ac1 temperature increases directly with heating rate and ranges from 706°C with a 1°C/s heating rate up to 759°C with
a 1000°C/s heating rate. The Ac3 temperature exhibits a much narrower range, 824° to 839°C, indicating that the ferriteto-austenite transformation finish temperature is not as dependent on heating rate as the ferrite-to-austenite transformation start temperature. The transformation start temperatures increase directly with the heating rate since the ferrite to austenite transformation is diffusion controlled. Increased heating rates provide less time for diffusion to occur, which causes a concomitant delay in the transformation start temperatures. These results were used to select peak temperatures to simulate the four critical regions of the HAZ. A peak temperature of 675°C was selected for the subcritical HAZ region since it is below the Ac1 temperature over the entire range of heating rates in vestigated. An 800°C peak temperature was selected for the intercritical HAZ region since it falls inside the Ac1 and Ac3 range for all four dilatometry curves. A 900°C peak temperature was selected for the fine-grained austenite region since 900°C minimally exceeds the maximum measured Ac3 temperature of 839°C. Finally, a 1350°C peak temperature was selected for the coarse-grained austenite HAZ region since 1350°C significantly exceeds the Ac3 temperature but is still below the melting temperature of the alloy. The Smartweld Optimization and Analysis Routine (SOAR) software (Ref. 13) was used to determine thermal cycles associated with the various peak temperatures identified previously. An 85% transfer efficiency, representative of the GMAW process, was assumed in the calculations (Ref. 14). Weld thermal cycles for each peak temperature were estimated for both a low (1.5 kJ/mm) and a high (3.75 kJ/mm) heat input to determine the effect of heat input on HAZ mechanical properties. The weld thermal cycles associated with the 675°, 800°, 900°, and 1350°C peak temperatures are shown in Fig. 3A, B for
A
B
Fig. 6 — Microhardness data collected from NUCu-140 base metal and simulated HAZ samples.
nificant austenite grain coarsening. This can be attributed to the short time and slight increase in temperature above Ac3. The presence of NbC particles also aids in restricting austenite grain growth. The microstructure of the LH1350 sample consists of acicular ferrite and a combination of either Widmanstätten ferrite, bainite, or low Fig. 5 — LOM micrographs of the following simulated samples: A — Low carbon martensite. heat input; B — high heat input simulated HAZ samples. The presence of acicular ferrite in the LH1350 sample indicates that significant the low and high input matrix, respecgrain coarsening occurred during the simtively. Each thermal cycle was subseulated HAZ thermal cycle since acicular quently linearized and input directly into ferrite nucleates intragranularly on hetthe Gleeble 3500 QuikSim software to erogeneous nucleation sites such as oxide produce simulated HAZ samples. inclusions, and its formation is enhanced Figure 4 shows the NUCu-140 base when the austenite grain size increases metal microstructure while Fig. 5A, B (Refs. 15–18). The grain coarsening also shows the microstructure of the low heat suggests that a 1350°C peak temperature input and high heat input simulated HAZ is high enough to dissolve the NbC partisamples, respectively. The low heat input, cles. This is consistent with thermody675°C (LH675) peak temperature sample namic calculations performed on similar exhibits a predominantly equiaxed ferritic Fe-Cu steels that indicate the NbC partimicrostructure that is nearly identical to cles will dissolve between approximately the base metal microstructure. Both the 1050° and 1100°C (Ref. 7). The high heat LH800 and LH900 samples also exhibit an input samples exhibit very similar miequiaxed ferritic microstructure with little crostructures when compared to the low to no change from the base metal miheat input counterparts, where the crostructure. Each of these simulated HH675, HH800, and HH900 samples all HAZ microstructures has a similar grain contain equiaxed ferrite. The HH1350 size to the base material, which indicates sample is also similar to its LH1350 counthat no significant grain coarsening octerpart in that it exhibits austenite grain curred as a result of the thermal cycle. The coarsening leading to an acicular-type miLH900 sample is fully austenitized during crostructure that contains a mixture of acithe thermal cycle, but does not exhibit sig-
cular ferrite, Widmanstätten ferrite, bainite, and low-carbon martensite. Figure 6 shows the average microhardness values for both the high and low heat input simulated HAZ samples. The low heat input results show a noticeable reduction in hardness for the LH675 (280 HV), LH800 (235 HV), and LH900 (225 HV) samples as compared to the base metal (300 HV). The observed softening in these simulated HAZ samples occurs as a result of evolution of the Cu-rich precipitates (Ref. 12). The LEAP tomography data shown in Fig. 1 exhibit a linear decrease in both the average precipitate radius () and precipitate volume fraction ( φ) as the weld interface is approached (i.e., increasing peak temperature). This is consistent with the reduction in hardness that is observed in the low heat input matrix. The LH1350 sample (290 HV) exhibits only a minimal reduction in microhardness even though it is expected to have undergone full dissolution of the Cu-rich precipitates. The relatively high hardness of the LH1350 sample is attributed to the acicular-type microstructure (mixture of acicular ferrite, Widmanstätten ferrite, bainite, and low-carbon martensite) observed in this region. The high heat input microhardness results exhibit a similar trend where there is local softening in each of the four simulated HAZ samples. The slightly lower hardness values observed for the high heat input 675°, 800°, and 900°C peak temperatures, relative to the low heat input samples, are probably the result of increased coarsening/dissolution that is associated with the longer heating and cooling times of the high heat input thermal cycle. The HH1350 sample (260 HV) again shows a slight hardness recovery as compared to the observed minimums that occurred in
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A
B
Fig. 8 — Charpy data collected for NUCu-140 at –40°C.
C
W E L D I N G R E Fig. 7 — Tensile testing results for both low and high input samples showing the following : A — S heat Yield; B — tensile; C — elongation. E A R the HH800 and HH900 samples. C Figure 7 shows the yield strength, tenstrength, and elongation results for the H sile high and low heat input samples. Each simulated HAZ data point presented in Figure 7 was generated from a sample that failed within the uniformly heated region. This was verified through direct temperature measurements during the simulation cycle and postfracture analysis. The low heat input yield and tensile strength values decrease as the HAZ thermal cycle peak temperature increases up to the LH900 sample. Partial recovery of the tensile
properties is observed in the LH1350 sample, which agrees with the observed microhardness trends. These strength trends are consistent with the observed microhardness results discussed previously. The elongation values only range from 18 to 21% across the base metal and all four simulated HAZ samples, with the ductility decreasing in the 675° and 800°C samples. This observation is unexpected since the 675° and 800°C samples have decreased strength levels compared to the base metal, which would typically lead to higher ductility. However, it must be noted that the specimen gauge length contains NUCu-140 material that was heated to a range of different peak temperatures during the Gleeble thermal cycle. The uniformly heated region in the simulated HAZ sample is shorter than the specimen gauge length and so the elongation results represent some average elongation behavior of each peak temperature/microstructural region. As a result, the elongation value reported for each condition is actually a composite measurement and the resulting trends are insignificant. Nearly identical trends in yield strength and tensile strength are observed for the high heat input sample matrix. Figure 8 shows the Charpy impact values for both the low and high heat input
samples tested at –40°C. The Charpy impact energy generally increases relative to the base metal value for the 675°, 800°, and 900°C peak temperature samples. This is consistent with the microhardness and yield strength results, since a decrease in strength/hardness typically produces an increase in toughness. The high heat input 1350°C peak temperature sample exhibits only a very slight reduction in impact toughness relative to the base metal, while the impact toughness of the low heat input sample is higher than that of the base metal. The difference in impact toughness between the LH1350 and HH1350 samples can be attributed to the prior austenite grain size in each region. The LH1350 sample has a prior austenite grain size of 32 µm while the HH1350 sample has a prior austenite grain size of 42 µm. The reduced prior austenite grain size and corresponding increase in grain boundary area in the low heat input sample results in a concomitant increase in the impact toughness. It is interesting to note that all regions of the HAZ exhibit relatively good impact toughness relative to the base metal, regardless of the location or heat input. In order to better understand the observed mechanical property trends, the expected precipitate evolution in the base
Table 3 — Calculated Precipitate Parameters and Measured and Predicted Strength Change for LH800, LH900, HH800, and HH900 Samples
Base Metal LH800 LH900 HH800 HH900
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Average Precipitate Radius (nm) 4.25 0.56 0.55 0.60 0.60
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Precipitate Volume Fraction φ 0.0099 0.0042 0.0041 0.0051 0.0051
Interprecipitate Spacing L (nm)
Predicted Strength Change (MPa)
Measured Strength Change (MPa)
75.5 15.1 15.0 15.0 14.9
— –162 –185 –52 –53
— –165 –183 –176 –176
Fig. 10 — Measured vs. predicted strength change of 800° and 900°C simulated HAZ samples using MatCalc as input to the Russell-Brown model.
Fig. 9 — MatCalc kinetic simulation for the LH675 simulated HAZ sample.
metal and HAZ regions was modeled using MatCalc kinetic simulations (Refs. 12, 19–22). The MatCalc method estimates precipitate nucleation and growth kinetics in multicomponent and multiphase alloys and can deal with complex systems and precipitation sequences. The MatCalc method accounts for the compositional dependence of interfacial energy and chemical driving force, which pro vides a significant improvement to the classical nucleation, growth, and coarsening models when applied to the Fe-Cu system. The simulations show the expected evolution of the Cu-rich precipitates in terms of average precipitate radius , number density (Nv), and precipitate volume fraction ( φ). An example MatCalc simulation result for the LH675 sample is presented in Fig. 9. The LH675 sample begins to undergo partial dissolution of the Cu-rich precipitates at the end of the heating portion of the HAZ thermal cycle as evidenced by the reduction in average precipitate radius and precipitate vol-
ume fraction (φ). The excess solute from the recently dissolved precipitates causes re-precipitation of new, smaller Cu-rich precipitates during the cooling portion of the cycle, which leads to a resultant increase in the precipitate number density. MatCalc simulations were also produced for the LH800, LH900, and LH1350 samples and in all three cases the Cu-rich precipitates fully dissolved on heating, followed by re-precipitation of the Cu-rich precipitates on cooling. An identical set of MatCalc simulations was performed for the high heat input samples and similar trends were observed. A summary of the predicted precipitate parameters for the base metal and each of the simulated HAZ samples can be seen in Table 2. Comparison of the base metal precipitate parameters generated using MatCalc and LEAP tomography reveals that generally good agreement is achieved for the and Nv. However, the MatCalc predicted φ (~0.0099) is more than three times less than the φ measured using LEAP tomography (~ 0.032). Previous work by the authors (Ref. 12) demonstrated that the significantly higher φ values measured using LEAP tomography can be explained by uncertainty in determining the true Cu concentration of the Cu-rich precipitates. A combination of empirical Cu solubility data, published precipitate Cu concentrations, and lever law calculations demonstrated that for the binary Fe-Cu system the expected φ of Cu-rich precipitates for NUCu-140 is approximately 0.013, which is in very good agreement with the MatCalc prediction of 0.0099. Based on this analysis (Ref. 12), it was decided that the MatCalc precipitate parameters would be used as input for the strengthening calculations (discussed below). Using the predicted precipitate parameters generated using MatCalc, the strengthening model proposed by Russell and Brown (Ref. 23) can be used to esti-
mate the change in strength expected from changes in the Cu-rich precipitates. The Russell-Brown model assumes a random distribution of spherical precipitates that are elastically softer than the surrounding matrix. These two assumptions are generally acceptable for NUCu-140 since the Cu-rich precipitates are roughly spherical and are elastically softer than the nominally ferritic matrix (Ref. 23). The RussellBrown model defines the shear strength of the material as the following: 1
2 2 Gb E P τ = 0.8 1 − L E 2 M when sin–1
E
p
≤ 50 deg
E
(1)
M
3
2 4 Gb E P τ = 1 − L E 2 M when sin–1
E
p
E
d ≥ 50 deg
(2)
M
E
log
P
E
P
E
M
r
∞
0
=
E
∞
log
M
R
r
0
L
=
R
log
r
r
+
(3)
R
log
r
0
1.77r
(4)
φ
where G is the shear modulus of the matrix (77 GPa); b is the Burgers vector (0.25 nm); L is the interprecipitate spacing (Equation 4 ); E P is the dislocation energy in the precipitate; E M is the dislocation energy in the matrix; E P ∞ is the dislocation WELDING JOURNAL
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A
B
11 — Sensitivity analysis of the Russell-Brown W Fig. model showing predicted increase in yield strength of radius for constant volume fracE astions.a function A — Full scale; B — enlarged view. L D in the precipitate per unit length; I N Eenergy is the dislocation energy in the matrix M G per unit length; r is the radius; r 0 is the inner cut-off radius (2.5 b); R is the outer R cut-off radius (1000 r 0); and φ is the preE cipitate volume fraction. The shear moduS lus of body-centered cubic (bcc) Cu is asto be equivalent to that of E sumed face-centered cubic (fcc) Cu. The ratio of A dislocation energy per unit length in the R precipitate and matrix ( E P )⁄( E M ) is apC proximately 0.6 (Ref. 23). The shear in each simulated HAZ region is H strength calculated by substituting in the precipi∞
∞
∞
tate parameters predicted using the MatCalc simulations (Table 2) and applying a Schmid factor of 2.5 (Ref. 23). The predicted shear strength change is then calculated by comparing the calculated base metal strength to the calculated strength in each simulated HAZ region. Application of the Russell-Brown model does have several limitations which include the assumption that the shear modulus of fcc Cu and bcc Cu have the same value, the calculation of L based on radius and precipitate volume fraction is subject to significant error (Ref. 23), and the cut-off radius values are geometric approximations. However, recent research has demonstrated that the model provides good agreement between calculation and experiment for similar Fe-Cu based alloys (Ref. 24). This calculation is assumed to be a valid indicator of the overall strength change since the matrix microstructure of the base metal, subcritical HAZ, intercritical HAZ, and fine-grained HAZ re-
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gions are composed of an equiaxed ferritic microstructure with nearly identical grain sizes. The coarse-grained HAZ region therefore cannot be compared using this method due to the acicular-type microstructure that results from the HAZ thermal cycle. Figure 10 shows the measured vs. predicted yield strength change for the 800° and 900°C peak temperature simulated HAZ samples for both the low and high heat input. The calculated shear strength values were converted into yield strength values by applying the relationship that shear yield strength is approximately equal to 0.6 times the tensile yield strength in steel (Ref. 25). The 1350°C samples were not included in the strength calculations since they exhibit acicular-type microstructures and therefore cannot be directly compared to the equiaxed ferritic structures. The 675°C peak temperature samples were included in the calculation but were not included in Fig. 10 since the strengthening model predicts an increase in yield strength of greater than 160 MPa for the low heat input sample, while the measured yield strength exhibits a decrease of approximately 60 MPa. This de viation can be explained as a result of the partial dissolution and re-precipitation predicted for the 675°C peak temperature samples (Fig. 9), as well as how the interparticle spacing (L) is calculated. L is calculated as a function of the average precipitate radius () and the precipitate volume fraction ( φ) through Equation 4, but does not account for the precipitate number density directly. This is typically not a problem since all three of these precipitate parameters are interrelated. However, for the case of the 675°C peak temperature simulated HAZ sample, re-precipitation of new, smaller Cu-rich precipitates is predicted during cooling. This leads to a bimodal precipitate distribution where the new, smaller precipitates are not correctly accounted for in the calculation of L. The new precipitates cause to decrease significantly, which results in a calculated L value that is too low and which leads to a predicted strength value that is too high. The measured vs. predicted strength change for the LH800 and LH900 samples exhibit very good quantitative agreement where there is an almost one-to-one relationship between the measured and predicted values. However, the quantitative agreement breaks down for the high heat input samples. The calculations for the high heat input condition accurately reflect that there is little difference between the HH800 and HH900 strength levels, which was observed experimentally in Fig. 7A, but the magnitude of the measured and predicted strength changes show significant disagreement. The predicted de-
crease in strength is approximately 50 MPa as opposed to nearly 175 MPa for the measured values. The predicted precipitate parameters and strength changes are shown in Table 3 for each heat input. Note that MatCalc predicts only slight differences in the precipitate parameters for the low and high heat input samples. This agrees with the experimentally measured strength levels, which show only minimal differences in the resultant strength. This suggests that MatCalc is accurately capturing the precipitate evolution for the various HAZ thermal cycles. This also implies that NUCu-140 is not very sensitive to heat input over the range of heat inputs evaluated in this study. The large variation in predicted strength calculated with only minor differences in and φ suggest that the Russell-Brown model is very sensitive to the input parameters of and φ. Figure 11A, B shows a sensitivity analysis that plots the increase in yield strength predicted by the Russell-Brown model as a function of for three representative φ values. The yield strength initially increases with increasing precipitate radius for a fixed volume fraction (Fig. 11A) up to approximately 1 nm, after which the yield strength is expected to decrease (Ref. 26). This is consistent with a coarsening phenomenon where the overall precipitate strength contribution decreases when the precipitates coarsen past the peak aged condition. Figure 11B shows an enlarged view of Fig. 11A for precipitate radii up to 1 nm. The MatCalc-predicted low heat input and high heat input precipitate parameters are labeled along with the associated strengthening contribution predicted by the Russell-Brown model. The slope of the strength vs. curve in the vicinity of the low heat input and high heat input precipitate parameters are 1590 and 2275 MPa/nm, respectively. Note that, within this region, very large strength increases are predicted with only minor increases in precipitate radii for this model. MatCalc simulations or LEAP tomographic measurements are probably not accurate enough to expect valid quantitative comparisons with the R-B model at the low end of the range when such small variations in precipitate parameters can cause large differences in the predicted strength. This effect probably accounts for the disagreement between the measured and calculated strength changes shown in Fig. 10. Conclusions
Microstructural evolution and mechanical properties of simulated heataffected zones in NUCu-140 steel was in vestig ated via light optical microscopy, dilatometry, Gleeble HAZ simulations, mechanical testing, and modeling tech-
niques. The following conclusions can be drawn from this research. 1. Dilatometry experiments over a wide range of heating rates show that the Ac1 temperature ranges from 706°C with a 1°C/s heating rate up to 759°C with a 1000°C/s heating rate. The Ac3 temperature exhibits a narrower range of 824° to 839°C. 2. The subcritical, intercritical, and fine-grained HAZ regions exhibit an equiaxed ferritic microstructure that is very similar to the NUCu-140 base metal microstructure. The coarse-grained HAZ exhibits a predominantly acicular-type matrix microstructure composed of a combination of acicular ferrite, Widmanstätten ferrite, bainite, and low-carbon martensite. 3. Microhardness and tensile testing results demonstrate that local softening occurs in the HAZ, with the minimum strength and hardness occurring in the intercritical and fine-grained HAZ regions. The strength loss in these regions is attributed to complete dissolution and only partial re-precipitation of Cu precipitates, as shown by precipitate simulation. The coarse-grained HAZ exhibits a slight recovery of strength and hardness as a result of the acicular-type structure that forms in this region. 4. The Charpy impact energy at –40°C for each HAZ region was equal to or better than the unaffected base metal. This is consistent with the microhardness and yield strength results, since the HAZ regions exhibited a decrease in strength/hardness, which typically produces an increase in toughness. 5. Similar values of precipitate radii, number density, and phase fraction were calculated for the 675°, 800°, and 900°C peak temperature samples for both low and high heat input conditions. The simi-
lar precipitate parameters and equiaxed ferrite microstructure would be expected to produce similar strength levels for each of these samples. This is consistent with the measured mechanical properties where there are only minor variations in the yield strength for these conditions as a function of heat input. This indicates that NUCu-140 is insensitive to heat input over the range of heat inputs (1.5–3.75 kJ/mm) investigated in this study. Acknowledgments
The authors gratefully acknowledge financial support of this research by the Office of Naval Research through Grant Number N00014-07-1-0331 and useful discussions with the program manager, Dr. William Mullins, of the Office of Naval Research. References
1. Isheim, D., and Seidman, D. N. 2004. Sur face and Interface Analysis 36: 569. 2. Isheim, D., Gagliano, M. S., Fine, M. E., and Seidman, D. N. 2006. Acta Materialia 54: 841. 3. Kolli, R. P., and Seidman, D. N. 2007. Mi croscopy and Microanalysis 13: 272. 4. Gagliano, M. S., and Fine, M. E. 2004. Metallurgical and Materials Transactions A 35A: 2323. 5. Isheim, D., Kolli, R. P., Fine, M. E., and Seidman, D. N. 2006. Scripta Materialia 55: 35. 6. Kolli, R. P., and Seidman, D. N. 2008. Acta Materialia 56: 2073. 7. Kolli, R. P., Wojes, R. M., Zaucha, S., and Seidman, D. N. 2008. International Journal for Materials Research (formerly Zeitschrift fur Met allkunde) 99: 513. 8. Kolli, R. P., and Seidman, D. N. 2011. International Journal for Materials Research (formerly Zeitschrift fur Metallkunde). 9. Leister, B. M., and DuPont, J. N. 2012. Fracture toughness of simulated heat-affected
zones in NUCu-140 steel. Welding Journal 91(2): 53-s to 58-s. 10. Fine, M. E., Ramanathan, R., Vaynman, S., and Bhat, S. P. 1993. International Sympo sium on Low Carbon Steels for the ‘90s, p. 511. 11. Bhadeshia, H. K. D. H., and Honeycombe, R. W. K. 2006. Steels: Microstructure and Properties. Elsevier, p. 287. 12. Farren, J. D., Hunter, A. H., DuPont, J. N., Seidman, D. N., Robino, C. V., and Kozeschnik, E. 2012. Submitted to Metallurgical and Materials Transactions A. 13. Fuerschbach, P. W., and Eisler, G. R. 2002. 6th International Trends in Welding Re search Conference Proceedings. 14. DuPont, J. N., and Marder, A. R. 1995. Thermal efficiency of arc welding processes. Welding Journal 749(12): 406-s. 15. Babu, S. S. 2004. Current Opinion in Solid State and Materials Science 8: 267. 16.Babu, S. S., and Bhadeshia, H. 1992. Materials Science and Engineering A A156: 1. 17. Babu, S. S., and Bhadeshia, H. 1990 Materials Science and Technology 6: 1005. 18. Babu, S. S., and David, S. A. 2002. ISIJ International 42: 1344. 19. Holzer, I., and Kozeschnik, E. 2010. Mater. Sci. Forum 638–642: 2579. 20. Kozeschnik, E., Svoboda, J., Fratzl, P., and Fischer, F. D. 2004. Materials Science and Engineering A — Structural Materials Properties, Microstructure and Processing 385: 157. 21. Kozeschnik, E. 2008. Scripta Materialia 59: 1018. 22. Svoboda, J., Fischer, F. D., Fratzl, P., and Kozeschnik, E. 2004. Materials Science and En gineering A — Structura l Materia ls Propert ies, Microstructure and Processing 385: 166. 23. Russell, K. C., and Brown, L. M. 1972. Acta Metallurgica 20: 969. 24. Yu, X., Caron, J. L., Babu, S. S., Lippold, J. C., Isheim, D., and Seidman, D. N. 2010. Acta Mater. 58: 5596. 25. Deutschman, A. D., Michels, W. J., and Wilson, C. E. 1975. Machine Design: Theory and Practice. 26. Kolli, R. P., and Seidman, D. N. 2011. International Journal for Materials Research (formerly Zeitschrift fur Metallkunde) 102: 1115.
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Preliminary Investigation on UltrasonicAssisted Brazing of Titanium and Titanium/Stainless Steel Joints Aluminum-based brazing filler metal can be used for brazing titanium to steel in air BY A. ELREFAEY, L. WOJARSKI, J. PFEIFFER, AND W. TILLMANN
ABSTRACT Preliminary investigations of the microstructure and fracture behavior of ultrasonicassisted brazing of CP titanium to itself and to AISI 304 stainless steel was conducted, using an aluminum-based filler. Test joints were processed at a temperature of 670°C and with a holding time of 3 min, followed by ultrasonic vibrations for 6 s. The resultant joints were characterized in order to determine the brittle intermetallic compound (IMC) in the interfacial layer. The shear strength of the joints was tested as well. The preliminary experimental results showed that sound joints with good wetting quality, without pores and cracks can be achieved. Intermetallic Ti-Al phases were detected at the titanium/aluminum-based filler metal in both similar and dissimilar joints. Both joints fractured after shear strength tests in the area containing this intermetallic compound. The titanium/titanium joints achieved a higher shear strength of 64 MPa. Meanwhile, the titanium/stainless steel joint obtained 46 MPa.
W E L D I N G Introduction R Titanium and its alloys exhibit a unique E combination of mechanical and physical S properties as well as corrosion resistance, E which make them desirable for several inA dustrial sectors such as power generation, processing, aerospace, and medR chemical ical applications. On the other hand, steel C and steel alloys represent the most imporH tant and widely used materials in industrial applications. Gradually, composite structures of dissimilar metals were accepted in national defense and civil industrial fields, such as aeronautics and astronautics, and energy and electric power industries. Composite components of titanium alloy and steel can take advantage of these two materials simultaneously. A partial replacement of steel components with titanium alloys will become an important way to reduce the mass of spacecrafts (Refs. 1−3). Titanium belongs to a family of metals called reactive metals that have a strong affinity for oxygen. At room temperature, titanium reacts with oxygen to form titanium dioxide. This passive, impervious coating resists further interactions with the surrounding atmosphere, and gives titanium its famous corrosion resistance. The
A. ELREFAEY, L. WOJA RSKI ( lukas.wo [email protected]) , J. PFEIFFER, and W. TILL MANN are with the Institute of Materials Engi neering, TU Dortman, Dortman, Germany.
148-s MAY 2013, VOL. 92
resultant layer has to be removed prior to joining because it melts at a much higher temperature than the base metal (Ref. 4). Brazing is one of the most inexpensive and convenient methods for joining titanium and titanium to dissimilar metals. Joining in a vacuum environment increases the cost of production and also reduces the design flexibility (Refs. 5, 6). Therefore, vacuum-free joining processes, which are assisted by external mechanical energy to disrupt the oxide layer in air, have been developed and investigated. Ultrasonic waves have been applied for soldering and brazing aluminum and titanium in air (Refs. 10–12). Ultrasonic vibrations imposed on metal surfaces cause a high cavitation intensity in the liquid filler metal, which disrupts and flakes off surface oxides, thereby allowing the filler metal to wet the surfaces and form a metallurgical bond. When undermining phenomena occurred during the interaction,
KEYWORDS Titanium Stainless Steel Brazing Ultrasonic Joint Microstructure Shear Strength
the oxide layer at the interaction interface was first lifted up by the undermining alloy, suspending in the liquid filler metal, before being broken up by the ultrasonic excitation (Ref. 10). In the case of nonultrasonic-assisted brazing of titanium, a stable oxide film grows on the surface of titanium alloys when heated in air. The oxide film presents a barrier for important interactions during the brazing process and resulted in incomplete wetting at the titanium interface. The filler metals mostly used for brazed titanium are silver-based (Refs. 11–15), titanium-based (Refs. 16–18), and aluminum-based filler metals (Refs. 19–21). Aluminum-based filler metals have the potential to braze titanium with sufficient properties. Their melting temperature ranges substantially below the beta transus, which make them a strong competitor to other filler metals. Other useful characteristics include lower densities and a good metallurgical compatibility with titanium alloys to be brazed, particularly good wetting and flow in capillary gaps. On the other hand, a lower shear strength of titanium brazed joints using aluminum filler metals can be generated by increasing the overlap area in order to achieve a load-carrying capability close to that of the base metal (Ref. 22). The main purpose of this study is to explore and evaluate the preliminary experiments related to brazing commercially pure titanium (CP Ti) and Cp Ti to stainless steel by using an aluminum-based filler metal (Al2.5Mg-0.3Cr) in open air utilizing an ultrasonic-assisted induction heating system. The focus is hereby on the interfacial microstructure and strengths of the joints.
Experimental Work The base metals used in this work were 2-mm-thick commercially pure titanium (CP Ti) Grade 2 and 2-mm-thick austenitic stainless steel AISI 304. The chemical compositions of the base metals are presented in Table 1. The stainless steel plate was cut into 25 × 25 mm chips,
and the titanium plate was cut into 10 × 10 mm chips, for shear strength testing and microstructure analyses. Additionally, the samples were first polished with SiC papers up to 1000 grit and subsequently cleaned by an ultrasonic bath using acetone as a solvent, prior to the brazing process. The filler metal used was a 50 µm thick TiBrazeAl-665A (Al-2.5Mg-0.3Cr, wt-%) designed for brazing thin-walled titanium articles and titanium matrix composites. This filler metal has a liquidus of 650°C, which is significantly lower than the beta transus temperature of CP Ti. The brazing foils, cleaned in acetone before brazing, were sandwiched between the overlapping areas of the base metals. The ultrasonic-assisted brazing apparatus used in this study is schematically illustrated in Fig. 1. A horn was installed in the vertical direction, and the test pieces were mounted into a steel holder. Initially, the samples were heated up to a temperature 50 K below the solidus temperature of the filler metal for a dwell time of 5 min, using a high-frequency induction coil in air. This step aimed at achieving the thermal equilibrium of the couple. The induction heating system has an output power of 15 kW and operating frequency of 13 KHz. The sample was then continuously heated up to 670°C with a holding time at this temperature of 3 min. The specimen temperature was measured by a K-type thermocouple, installed to touch a groove in the sample as close as possible to the joining interfaces. The horn was kept outside the heating region without preheating till the temperature of the samples was reached, and then it was moved manually to touch the surface of the sample before it started to work. Ultrasonic vibration with 120-W power at a frequency of 25 kHz was applied for 6 s and propagated in a direction perpendicular to brazing surfaces. This duration was recommended to completely destroy the oxide layer (Ref. 10). The samples were subjected to an equal pressure of nearly 0.2 MPa, which was a result of the weight of the horn. The average heating rate was 2.3 K/s, and the samples were cooled in air to room temperature after brazing. Selected samples were cut, mounted, polished, and etched for microscopic evaluation. A light optical microscope and scanning electron microscope (SEM), equipped with an energy-dispersive spectrometer (EDS), were used to characterize the joints. A metallographic examination was carried out on the cross section. Hardness measurements were performed with the help of a Vickers hardness testing machine with 25-g load and 25-s impressing time. Additionally, lap shear tests were performed to evaluate the bonding strength of the specimen as schematically illustrated in Fig. 2. The test
Fig. 1 — Schematic illustration of the ultrasonic assisted brazing apparatus.
was carried out at room temperature, and the displacement speed was 0.1 mm/s. Three samples were used to calculate the average shear strength. After the shear test, the fracture surfaces were analyzed by SEM and EDS analyses.
Results and Discussion The characteristic microstructures of the base metals are shown in Fig. 3. The microstructure of the stainless steel is composed of austenitic equiaxed grains with annealing twins in the grain interiors. Inclusions of the size of several micrometers can also be detected, while the microstructure of CP Ti possesses equiaxed α-phase grains.
Fig. 2 — Schematic illustration of the lap shear test.
Brazing of Titanium/Al-2.5Mg-0.3Cr/ Titanium Joint
Figure 4A displays SEM microstructure features of the joint. Obviously a sound joint was obtained since a homogeneous microstructure without voids or cracks was observed along the joint. The titanium showed no change in the microstructure since the brazing temperature is much lower than the β-Ti transformation temperature. According to the Ti-Al binary phase diagram (Ref. 23), the solubility of Al in α Ti at the brazing temperature is about 8 wt-%. Therefore, the titanium base metal was enriched by aluminum. The EDX analyses of the titanium base metal, close to the interfacial-brazed
Table 1 — Chemical Composition of Base Metals
Materials
CP Ti AISI 304
Wt-% C
Fe
Ti
Cr
Ni
Si
P
N
H
O
0.02 0.06
0.03 Bal.
Bal. —
— 17.88
— 8.52
— 0.31
— 0.009
0.03
0.01 —
0.25 —
Table 2 — Chemical Analyses at Areas Shown in Fig. 4
Table 3 — Chemical Analyses at Areas Shown in Fig. 7
Area Average Chemical Analyses (at.-%) Ti Al Mg Cr
Area Average Chemical Analyses (at.-%) Ti Al Mg Cr
1 2 3 4 5 6 7
98.76 99.61 0.37 74.60 61.46 68.70 68.63
1.24 0.39 97.35 21.99 36.58 30.12 29.96
— — 2.17 2.41 1.87 1.18 1.30
— — 0.11 — 0.09 — 0.11
1 2 3 4 5 6 7 8
5.55 6.33 97.09 92.51 69.35 69.97 66.70 61.06
91.36 90.71 2.51 6.24 27.53 26.42 23.75 32.09
2.96 2.85 0.40 1.25 3.04 3.49 9.48 6.83
0.13 0.10 — — 0.08 0.12 0.07 0.02
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A
B
B
A
Fig. 3 — Microstructures of base metals. A — Stainless steel AISI 304; B — CP titanium.
Fig. 4 — SEM microstruc ture features of the titanium/tita nium joint. A — General view of the cross section; B — close-up view at the interfacial area.
W E L D I N Fig. 5 — Hardness distribution in the titanium/titanium joint. Fig. 6 — Fracture path of the titanium/titanium joint. G R E area, contained a considerable amount of the layer indicates that it is almost a Ti 3 Al measured as well. In spite of the formation S titanium as shown in Fig. 4A, Areas 1 and phase — Area 4. This layer has a maxi- of the hard and brittle Ti 3 Al phase at the Table 2 lists the chemical analyses of all mum thickness of less than 2 m and is interfacial brazed area, the presence of E 2.phases at the brazed joint. The brazed formed discontinuously at the interface. this phase is preferred to as the TiAl phase A area mainly consists of a solid solution of Owing to the low ductility and toughness Ti 3 Al has much better strength and R aluminum dissolving a few percentages of of Ti3 Al at room temperature, this phase is since ductility than TiAl (Ref. 24). The hardness distribution in the joint is C magnesium, titanium, and chrome (Area continuously crushed during sample 3). It is worth noting that there was no inpreparations for metallographic investigashown in Fig. 5. The hardness of the H dication of an oxide layer at the interfacial tions, and hence, scattered to the soft alu- brazed area showed lower values than in µ
area. Ultrasonic vibration disrupted and flaked off the surface oxides as was explained in detail by Xu et al. (Ref. 10). The titanium/aluminum-based brazing interface is planar in nature, and a thin interaction layer was revealed at the interface area as shown in Fig. 4B. The stoichiometric ratio between Al and Ti of
minum brazed area. The scattered phase .can be easily detected in the brazed area as clearly shown in Fig. 4A, Areas 5 and 6, and the close-up view in Fig. 4B, Area 7. Energy dispersive X-ray (EDX) analyses of the scattered Ti 3 Al reflected more aluminum content since the phase is too thin, and its aluminum background area was
Table 4 — Chemical Analyses at Areas Shown in Fig. 8
Area 1 2 3 4 5 6 7 8 9
150-s
Average Chemical Analyses (at-%) Ti 70.19 20.19 73.72 1.94 30.19 0.19 0.99 99.53 0.17
Al 22.93 73.98 19.28 83.91 63.98 86.73 80.31 0.47 2.71
MAY 2013, VOL. 92
Mg 6.00 3.58 3.85 4.91 3.58 — 0.65 — 0.20
Cr 0.36 0.67 1.15 1.31 0.67 5.51 3.19 — 19.20
Fe 0.43 0.32 0.82 7.17 0.32 5.26 13.58 — 68.73
Ni 0.09 0.06 0.08 0.76 0.06 2.31 1.28 — 7.77
Si — 1.20 1.10 — 1.20 — — — 1.22
the titanium base metal. Additionally, the hardness close to the interfacial area in the titanium side was higher than in the base metal far from the interface, owing to the diffusion of aluminum into the base metal. However, the hardness at the interfacial area, which is expected to present the highest values, could not be assessed due to the very thin interfacial area. The fracture shear strength was calculated as the failure load, divided by the overlap area. The achieved average shear strength of the joints is 64 MPa with a standard deviation of ±2.7 MPa. This result is comparable or a little lower than results from our previous work, related to brazing titanium in a vacuum, using a sil ver-based alloy (Ref. 15). The strength of the joint could be improved by optimizing the brazing parameters such as the brazing temperature, holding time, and ultrasonic time. Figure 6 shows that the joints failed mainly at the titanium/aluminumbased filler metal interface owing to the
formation of the hard and brittle Ti 3 Al intermetallic compound. The fracture morphology of the joints after the shear test is presented in Fig. 7A. Chemical analyses of the corresponding fracture area (Table 3) showed a high probability of a Ti3 Al phase at the surface of the aluminum-based filler metal — Areas 5 and 6 in Fig. 7B, and the titanium base metal as well, Areas 7 and 8 in Fig. 7C. This implies that the Ti 3 Al intermetallic compound is the most harmful phase in the joint. Aluminum-based filler metal is shown in the fracture surface by the Areas 1 and 2, while the titanium base metal is presented by the Areas 3 and 4. The fractography of these fracture surfaces basically showed cleavages in addition to tearing regions. The fracture direction took the same direction as in the shear test.
A
Brazing of Titanium/Al-2.5Mg-0.3Cr/ Stainless Steel Joint
Figure 8A displays SEM microstructure features of the joint. It is clear that the brazed filler metal wets the titanium and stainless steel base metals well, and no defects are observed at the interface. The brazed area is mainly composed of three zones. The first zone is the interfacial thin reaction layer at the titanium/aluminumbased filler metal. Energy-dispersive X-ray analyses of this layer showed that it is the Ti-Al intermetallic phase, either a Ti 3 Al or an Al3Ti or a mix of them — Fig. 8B, Areas 1–3. Table 4 lists the chemical analyses of areas shown in Fig. 8. It is difficult to exactly identify the content of this phase since it is very thin and fragmented at the interface. The second zone is close to the previous interaction layer and basically consists of aluminum-based matrix (Area 4) with fragments of Ti-Al intermetallic phases (Area 5) separated from the interfacial area similar to the titanium/Al2.5Mg-0.3Cr/titanium joint. The third zone is the one close to the stainless steel/aluminum-based filler metal. In this area, the solubility of iron into the aluminum is almost zero, which results in a very early interfacial phase formation when iron is dissolved into aluminum. Iron and chromium diffusion from the stainless steel side constitute several phases with the aluminum-based filler metal. The isothermal section in the Al-Fe-Cr ternary phase diagram at 600°C and the partial isothermal section at the aluminum-rich corner (more than 75 at.-% Al) confirm the existence of three-phase (1) υ +(2) θ + (Al) and υ + (Al13Fe4) + (Al) (25–28) — (1) The stoichiometric chemical composition of this phase is Cr 10.71 Fe8.68 Al80.61 , and it is sometimes called H-CrFeAl. (2) The stoichiometric chemical composition of this phase is Cr 2 Al13 and sometimes called CrAl 7 in the literature.
C
B
Fig. 7 — Fracture morphology of the titanium/titanium joint. A — General view of the fracture surface; B — close-up view at the surface of aluminum-based filler; C — close-up view at the surface of titanium base metal.
Fig. 9A and B, respectively. Energy-dispersive X-ray analyses at the interfacial area suggest the presence of the threephase υ + (Al13Fe4) + (Al) in spite of the low accuracy of the analyses due to the minute size of the phases. The three-phase are shown in Fig. 8C, Area 6 (grey grains of phase υ surrounded by black solid solution Al) and Area 7 (Al 13Fe4 intermetallic phase). Similar to the titanium/titanium joint, the diffusion of aluminum into the titanium base metal in the titanium/stainless steel joint was shown in Fig. 8, Area 8. On the other side of the joints, the solubility of Al in the austenitic matrix is on the order of about 2 to 2.5 wt-% Al. Therefore, EDX analyses detected a few percentages of aluminum in the stainless steel side of the joint — Fig. 8, Area 9. It is to be noted that the size and distribution of intermetallic compounds were different from the Ti/Ti joint to the Ti/stainless steel joint. The solubility of Ti in aluminum at brazing temperature is almost neglected (Ref. 23). Therefore, high
percentages of titanium caused the formation of Ti 3 Al phase only at the interface since titanium has limited solubility in molten aluminum. On the other hand, iron has a high solubility in molten aluminum and is easily dissolved at the molten stage (Refs. 29, 30). During cooling, iron has a very low solubility in the solid state and is therefore present mostly as a coarse intermetallic phase in the brazed zone. The hardness distribution in the joint is shown in Fig. 10. In contrast to the titanium/titanium joint, the hardness of the brazed area showed the highest values. The brazed area close to the stainless steel side showed the peak hardness in respect to other areas in the brazed zone. It was also noted that the stainless steel showed a higher average hardness than titanium. The average shear strength of the joints achieved 46 MPa with a standard deviation of ± 3.1 MPa. The strength was lower than the titanium/titanium joint, since an Al-Fe intermetallic compound was detected in the brazed area in addition to Al-Ti. Addi-
Table 5 — Chemical Analyses at Areas Shown in Fig. 11
Area 1 2 3 4 5 6
Average Chemical Analyses (at.-%) Ti 5.98 7.58 11.52 22.52 26.77 31.32
Al 85.69 84.52 75.57 56.55 53.49 61.01
Mg 6.16 3.30 6.91 18.45 16.77 7.55
Cr 0.79 0.59 0.82 0.53 0.78 0.12
Ni 0.25 0.31 0.52 — — —
Fe 1.13 3.70 4.66 1.95 2.19 —
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A
C
B
Fig. 8 — SEM microstructure features of the titanium/stainless steel joint. A — General view of the cross section; B — close-up view a t the titanium/aluminum-based filler metal interfacial area; C — close-up view at the stainless steel/aluminum-based filler metal interfacial area.
W A B E L D I N G R E S E A R Fig. 9 — A — Isothermal section in the Al-Fe-Cr ternary phase diagram at 600°C; B — the partial isotherC mal section at the aluminum-rich corner. H tionally, extra internal stresses are expected in this joint compared with the similar titanium joint. Surprisingly, Fig. 11 shows the joints failed mainly at the titanium/aluminum-based filler metal interface in spite of the thinner intermetallic compound of this area, compared with the area close to the stainless steel/aluminumbased filler metal, which showed thick and different intermetallic compounds. Fracture surface, corresponding to the previous fracture pass, is shown in Fig. 12A and C for the stainless steel and titanium sides, respectively. Chemical analyses of different areas at the fracture surface generally showed high aluminum content in the stainless steel side (Table 5, Areas 1–3). Meanwhile, at the titanium side, the content of titanium increased significantly (Areas 4–6). The stoichiometric composition of different areas did not confirm the occurrence of any Al-Ti intermetallic compound at the fracture surface in contrast to the titanium/titanium joint. 152-s
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Figure 12B and D showed enlarged images of the stainless steel and titanium sides, respectively. The microscope fractography of these fracture surfaces basically showed cleavage morphologies in both sides of the joint with more tearing regions in the stainless steel side of the fracture and more flat areas and shearing directions in the titanium side. Conclusions
Ultrasonic-assisted brazing experiments of CP titanium to itself and to AISI 304 stainless steel were conducted using an aluminum-based filler metal. The joints were successfully brazed without voids, cracks, or surface oxides disturbing the wetting of the joint. The relationship between the mechanical properties of the joints and the microstructure of the brazed layers was examined. The results obtained can be summarized as follows: 1) For the titanium/titanium joint, the
brazed area mainly consisted of solid solution aluminum with a Ti 3 Al intermetallic compound at the interfacial area. During shear tests, the crack pass propagated at this intermetallic compound has almost no ductility to withstand thermal stresses. The average shear strength of the joints was 64 MPa. 2) For the titanium/stainless steel joint, Ti-Al intermetallic compounds were formed at the titanium/aluminumbased filler metal interfacial area. Mean while, three-phase υ + (Al13Fe4) + (Al) were formed at the stainless steel/aluminum-based filler metal interfacial area. In spite of the high hardness of this area in respect to the titanium/aluminumbased filler metal interfacial area, the crack pass during shear tests was close to the Ti-Al intermetallic compound. The average shear strength of the joints was 46 MPa. References 1. Boyer, R. 1996. An overview on the use of titanium in the aerospace industry. Materials Science and Engineering A 213(1–2): 103–114. 2. Yuan, X. J., Sheng, G. M., and Qin, B. 2008. Impulse pressuring diffusion bonding of titanium alloy to stainless steel. Materials Char acterization 59(7): 930–936. 3. Wang, T., Zhang, B., and Chen, G. 2010. Electron beam welding of Ti-15-3 titanium alloy to 304 stainless steel with copper interlayer sheet. Transactions of Nonferrous Metals Society of China 20(10): 1829–1934. 4. Luck, J., and Fulcer, J. 2007. Titanium welding 101: Best GTA practices. Welding Jour nal 86(12): 26–31. 5. Liu, L. M., Zhu, M. L., Pan, L. X., and Wu, L. 2001. Studying of micro-bonding in diffusion welding joint for composite. Material s Science and Engineering A 315 (1–2): 103–107. 6. Shi, L., Yany, J., Han, Y., and Peng, B. 2011. Behaviors of oxide layer at interface between semi-solid filler metal and aluminum matrix composites during vibration. Journal of Materials Science Technology 27(8): 746–752. 7. Watanabe, T. 2000. Soldering of high strength aluminum alloys with the aid of ultrasonic vibration. Proc. Int. Brazing & So ldering Conf . pp. 523. Albuquerque, N.Mex. 8. Watanabe, T., Yanagiswa, A., Furkawa, A., and Onuma, S. 1993. Soldering of Al-Mg alloy with the aid of ultrasonic vibration. Quarterly Journal of Japan Welding Society 11(4): 484–489. 9. Zhao, W. W., Yan, J. C., Yang, W., and Yang, S. Q. 2008. Brazing of aluminium matrix composites. Science and Technology of Welding and Joining 13(1): 66–69. 10. Xu, Z. W., Yan, J. C., Zhang, B. Y., Kong, X. L., and Yang, S. Q. 2006. Behaviors of oxide film at the ultrasonic aided interaction interface of Zn-Al alloy and Al 2O3p /6061Al composites in air. Materials Science and Engineering A 415(1–2): 80–86. 11. Ma, Z., Zhao, W., Yan, J., and Li, D. 2011. Interfacial reaction of intermetallic compounds of ultrasonic-assisted brazed joints between dissimilar alloys of Ti6Al4V and Al4Cu1Mg. Ultrason Sonochemistry 18(5): 1062–1067.
Fig. 10 — Hardness distribution in the titanium/stainless steel joint.
12. Chan, H. Y., Liaw, D. W., and Shiue, R. K. 2004. Microstructural evolution of brazing Ti-6Al-4V and TZM using silver-based braze alloy. Materials Letters 58: 1141–1146. 13. Shiue, R. K., Wu, S. K., and Chen, S.Y. 2003. Infrared brazing of TiAl using Al-based brazed alloys. Intermetallics 11: 661–671. 14. Liaw, D. W., and Shiue, R. K. 2005. Brazing of Ti-6Al-4V and niobium using three sil ver-based alloys. Metallurg ical and Materials Transactions 36A(9): 2415–2427. 15. Elrefaey, A., and Tillmann, W. 2009. Effect of brazing parameters on microstructure and mechanical properties of titanium joints. Journal of Material s Processing Technology 209(10): 4842–4849. 16. Rabinkin, A., Liebermann, H., Pounds, S., Taylor, J., Reidinger, F., and Siu-Ching, L. 1991. Amorphous Ti-Zr-base Metglas®‚ brazing filler metals. Scripta Metallurgica 25(1): 399–404. 17. Botstein, O., Schwarzman, A., and Rabinkin, A. 1996. Induction brazing of Ti-6Al-4V alloy with amorphous 25Ti-25Zr-50Cu brazing filler metal. Materials Science and Engineering A 206(1): 14–23. 18. Knepper, P., and Lohwasser, D. 2001. High-temperature brazing of titanium structures. DVS-Berichte 208: 83–88. 19. Wells, R. R. 1975. Low-temperature large area brazing of damage tolerant titanium structures. Welding Journal 54(10): 348-s to 356-s. 20. Sohn, W. H., Bong, H. H., and Hong, S. H. 2003. Microstructure and bonding mechanism of Al/Ti joint using Al-10Si-1Mg filler metal. Material s Science and Engineering A 355(1–2): 231–240. 21. Shiue, R. K., Wu, S. K., and Chen, S. Y. 2003. Infrared brazing of TiAl using Al-based brazed alloys. Intermetallics 11: 661–671. 22. Shapiro, A. E., and Flom, Y. A. 2007. Brazing of titanium at temperatures below 800°C: Review and prospective applications. DVS-Berichte 243: 254–267. 23. Baker, H. 1992. Alloy Phase Diagrams. ASM Handbook, ASM International, Materials Park, Ohio. 24. Froes, F. H., Suryanarayana, C., and Eliezer, D. 1992. Review: Synthesis, properties
Fig. 11 — Fracture path of the titanium/stainless steel joint.
A
B
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C
D
Fig. 12 — Fracture morphology of the titanium/stainless steel joint. A — General view of the fracture sur face at the stainless steel side; B — enlarged view at the fracture surface at the stainless steel side; C — general view of the fracture surface at the titanium side; D — enlarged view at the fracture surface at the titanium side.
and applications of titanium aluminides. Jour nal of Materials Science 27(19): 5113–5140. 25. Materials Science International Team MSIT®, and Ghosh, Gautam, Korniyenko, Kostyantyn, Sidorko, Vladislav, Velikanova, Tamara: Al-Cr-Fe (Aluminium-ChromiumIron). Eds. G. Effenberg and S. Ilyenko. Springer Materials — The Landolt-Börnstein Database ( http://www.spring ermateria ls.com). DOI: 10.1007/10915943_30. 26. Mo, Z. M., Zhou, H. Y., and Kuo, K. H. 2000. Structure of υ-Al80.61Cr10.71Fe8.68, a giant hexagonal approximant of a quasicrystal determined by a combination of electron microscopy and X-ray diffraction. Acta Crystallo graphica, Section B 56(3): 392–401.
27. Cornish, L., Saltykov, P., Cacciamani, G., and Velikanova, T. 2003. Al-Cr (Aluminum – Chromium). MSIT Binary Evaluation Pr ogram, ed. Effenberg, G., MSIT Workplace, Materials Science International Services GmbH, Stuttgart, Germany. 28. Audier, M., Durand-Charre, M., Laclau, E., and Klein, H. 1995. Phase equilibria in the Al-Cr system. Journal of Alloys and Compounds 220 (1–2): 225–230. 29. Dybkov, V. I. 1990. Interaction of 18Crl0Ni stainless steel with liquid aluminum. Jour nal of Material Science 25: 3615 to 3633. 30. Kattner, U. R. 1990. Al-Fe (AluminumIron), Binary Alloy Phase Diagrams, 2nd edition, ed. T. B. Massalski 1: 147–149.
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Dynamic Control of the GTAW Process Using a Human Welder Response Model A model was implemented to adjust the welding current in response to the characteristic parameters of the 3D weld pool surface to maintain consistent, complete joint penetration in GTAW BY W. J. ZHANG AND Y. M. ZHANG
ABSTRACT
W E L D I N G R E S E A R C H
In the modern welding industry where automated welding tends to be the mainstream, manual welding is still not replaceable when human experience and skills are critical to produce quality welds. Yet the mechanization and transformation of a human welder’s intelligence into robotic welding have not been explored. In our previous study to understand a human welder’s behavior, the welder’s adjustments on welding current were modeled as a response to characteristic parameters of the three-dimensional weld pool surface. In this work, the response model is implemented to feedback control the gas tungsten arc welding (GTAW) process to maintain consistent, complete joint penetration. Experiments were designed to start welding using different welding conditions (arc length, welding speed, and root opening) along with initial current. After the initial openloop control period, the welding current is adjusted by the controller that uses the welder’s response model to determine how to adjust the welding current based on the measured weld pool surface characteristic parameters. The resultant current waveform and its backside weld bead width were recorded/measu red and analyzed. It was found that the human welder response model can adjust the current appropriately to control the welding process to a desired penetration level despite the difference in the welding conditions and initial current. The desired backside width of the weld bead, 5.2 mm, was produced with a 0.4 mm variation successfully in all experiments despite their diverse welding conditions and initial current.
Introduction Manual gas tungsten arc welding (GTAW) is thought by many as an operation that requires the highest skills, yet is commonly used in the industry, especially for applications requiring assured weld quality. A human welder can hear the sounds of the arc, sense the reactive forces from the torch, and observe the weld pool surfaces. Using such feedback information, a welder can appraise the welding process with respect to the desired state, then intelligently adjust the welding parameters (e.g., current, welding speed, arc length), and maintain appropriate torch orientation and distance in an effort to W. J. ZHANG and Y. M. ZHANG ([email protected]) are with the Institute for Sustainable Ma nufacturing and Department of Electrical and Computer Engineering, Univer sity of Kentucky, Lexington, Ky.
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control the desired weld state. Because of their experience-based behavior in response to the information they sense, human welders may be preferred over mechanized welding control systems in certain applications. Although welders’ experience and skills are crucial to producing quality welds, human welders have limitations. Critical welding operations require welders concentrate consistently to react
KEYWORDS Human Welder Response Joint Penetration Control Intelligent Welding and Control Weld Pool Surface Complete Joint Penetration Gas Tungsten Arc Welding (GTAW)
rapidly and accurately. Inconsistent concentration, fatigue, and stress build up such that welders’ capabilities degrade during daily operations. Moreover, experience and skills needed for critical operations typically require years to develop while the manufacturing industry is experiencing an insufficient number of skilled welders for a long time (Ref. 1). The mechanism of welders’ experience-based behavior, i.e., how welders respond to the information they acquire from their sensory system, should be explored and utilized to develop intelligent robotic welding systems that combine intelligence and physical capabilities for the next generation of manufacturing. Exploring the mechanism may also be utilized to understand why less skilled welders are not performing as well as skilled welders and help train welders faster to help resolve the skilled welder shortage issue the manufacturing industry is facing (Ref. 2). However, developing a model of the welders’ experience-based behavior and adapting it as a controller in automated welding is so far a challenging task. Numerous studies have been conducted with different sensing techniques mimicking welders’ sensing capability to the weld pool. Various types of information about the weld pool have been extracted and interpreted to describe the state of the welding process (Refs. 3–9). Although successes in monitoring the weld pool continue to be made in the academic community, the intelligent behavior of a human welder has not yet been successfully transferred to automated welding. This is because welders, in the role of human controller in the welding process, make decisions primarily based on past learned experiences, which might not in volve a fundamental understanding of the laws of physics. Also, a skilled welder assesses and controls a welding process using a humanistic approach where the feedback sensory information acquired by
the welder is imprecise and can only reflect partial truths about the instant status of the weld process. An automated welding control system requires both mechanistic methods for the welding phenomena that are physically well understood and mathematically feasible for both sensors and control algorithms. The theory of modeling for the human controller dynamics has been extensively studied since the 1940s. Great progress was achieved in the 1960s and 1970s (Ref. 10), such as linear crossover model (Ref. 11) and the optimal control model (Ref. 12). The physical nature of a human operator indicates that the human controller is naturally dynamic, stochastic, nonlinear, and time varying. In this sense, nonlinear methods were introduced to model the human action neural networks, and neuro-fuzzy or adaptive models (Refs. 13–17). Although nonlinear methods typically improve the prediction performance to some extent, it is still very appealing to use linear models due to their convenience for analysis and design. Instead of taking real industrial processes, most of the literature in this area took certain benchmarks as control objects, such as the pendulum, joystick, etc. Besides, those developed models tend to be too complex to understand and difficult to apply to the practical control systems. In our first study on human welder responses (Refs. 18, 19), dynamic models of a novice human welder’s behavior were developed. The studied behavior of the welder is focused on the adjustment of welding current in response to the observed three-dimensional (3D) weld pool surface during the complete-joint-penetration process. The weld pool geometry is used as the sensory feedback information since it is believed to provide valuable insights into the welding process state. Important information such as weld defects and penetration are contained in the surface deformation of the weld pool in the GTAW process (Refs. 20, 21). The
Fig. 1 — Demonstration of a manual control system of the GTAW proce ss. It is not a typical manual GTAW process. The human welder only adjusts the welding current based on his observation of the 3D weld pool surface. The pipe rotates during the experiment while the torch, imaging plane, laser, and cam era are stationary.
geometry of the weld pool has been studied (Refs. 22–26) as a means of monitoring and controlling the weld joint penetration. A vision-based sensing system has been developed to simultaneously measure the 3D weld pool surface and record the responses the human welder made to the surface. A dynamic model that correlates the welder responses (model outputs) to the characteristic parameters (model inputs) of the 3D weld pool surface has been established. This paper is the first of this kind addressing implementation of the human welder response model as a controller in the automated GTAW process. In particular, this study focuses on how this model controls the current to achieve consistent complete joint penetration under different welding parameters. The backside weld bead width is used as a measurement
for the penetration state. The effectiveness and robustness of the model-based control are evaluated and verified in this paper. Modeling of the human welder response is briefly reviewed in the next section. In the experimental system and methods section, a vision-based sensing system is detailed as well as the experiment method for implementation of the model. The results of the model-based control is presented and analyzed in the human welder response model control section. The human welder response model is further improved in the improvement of the human welder response model section. The robustness of the control using the improved model is then analyzed in the results and analysis of robustness experiments section. The conclusion is then given.
Table 1 — Experimental Parameters
Welding Parameters Root opening/mm/s [0, 5]
Arc length/mm [2, 5]
Welding speed mm/s 1.0
Initial welding current/A [50, 62]
Argon flow rate/L/min 11.8
Monitoring Parameters Project angle/deg 35.5
Laser to weld pool distance/mm 24.7
Imaging plane to tungsten axis distance/mm 101
Camera Parameters Shutter speed/ms 4
Frame rate/fps 30
Camera to imaging plane distance/mm 57.8
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12 o’clock without a filler metal. A human welder observes the weld pool and adjusts the welding current using an amperage remote control installed on the torch. The use of the remote controller for the welding current shown in the figure is for demonstration purposes only. The actual current remote controller is a thumb turn knob on the torch. It adjusts the current setting for the power supply. Vision-Based Sensing Subsystem
A
W E L D I N G R E S E A R C H
C
B
D Fig. 2 — Results of image processing and three dimensional reconstruction. A — Captured image using the sensing system; B — resultant dots in the captured image using image processing. The asterisk in the figure is the reference dot matching the dot a t the 10th ro w and 10th column in the pro jected laser dot matrix. C — Projected dots on the 3D weld pool surface; D — interpolated 3D weld pool surface; E — weld pool boundary and the pro jected dots in oxy plane. The pentagrams are the re flected laser dots, a nd the stars are the boundary dots of the weld pool. The blue curve is a fitted 2D weld pool boundary in literature (Ref. 28).
E
Human Welder Response Model Principle of Human Welderʼs Behavior
A skilled welder starts a welding process with initial welding parameters that are considered optimal based on past experiences. After observing the weld pool surface until enough feedback information is perceived, the welder assesses the process and adjusts the welding parameters accordingly to produce desirable welds. Skilled human welders are believed to make an optimal or nearly optimal control to minimize the error
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between the current and desired states of the welding process. Ideally, qualified/skilled welders make similar welds that meet the requirements because they all possess the ability to sense the process and make a decision using the sensed process feedback. Manual GTAW Experimental System
With the principle of the human welders’ behavior, an experimental system has been developed (Refs. 18, 19) as shown in Fig. 1. The pipe is rotated and butt joint welded using DCEN GTAW at
The 3D weld pool surface being observed by the human welder is also simultaneously measured by a vision system. The system includes the low-power, 20mW illumination laser generator at a wavelength of 685 nm with variable focus, a 19×19 dot matrix structured light pattern (Lasiris SNF-519X (0.77)-685-20) attached to the head of the laser, an imaging plane made by a piece of glass attached by a sheet of paper, and a camera (Point Grey Flea 3). The laser projects the 19 × 19 dot matrix on the melting region. Part of the dot matrix projected inside the weld pool is reflected by the specular weld pool surface. Then a reflection pattern of the dot matrix is intercepted by the imaging plane. Because of the plasma impact, the surface of the weld pool is depressed and distorted in GTAW. Therefore, no matter which shape (concave or convex) the weld pool presents, the alignment of the reflected laser dot matrix is distorted by the deformed specular weld pool surface. The distortion of the reflected dot matrix is determined by the shape of the threedimensional weld pool surface and contains the 3D geometry information about the weld pool surface. The camera captures the images of the reflected laser dot matrix from the imaging plane. A computer connected to the camera processes the images and reconstructs the 3D weld pool surface in real time (Ref. 27). Taking Fig. 2A, an acquired image in the imaging plane, as an example, the results of image processing and reconstruction are shown in Fig. 2B–E. The time for the image capturing, processing, and weld pool reconstruction is about 30 ms, which is fast enough for monitoring the weld pool dynamics in GTAW. Human Welder Response Model
Having the 3D weld pool surfaces recorded together with the current adjustment made by the human welder, the human welder response model establishes the correlation between the current adjustment and weld pool characteristics parameters, i.e., the length ( L), width ( W ), and convexity (C) of the 3D weld pool surface.
A
B
Fig. 3 — Weld pool boundary and parameters. A — 2D boundary; B — longitudinal intercepted area.
H C R A E S E R G N I D L E W
Table 2 —Welding Parameters Used in Experiments with Different Initial Currents
Welding current/A Arc length/mm
50, 54 3
Table 3 — Welding Parameters for Initial Current Robustness Experiments
Welding current/A Arc length/mm
50, 54, 58, 62 3
Table 4 — Welding Parameters for Arc Length Robustness Experiments
Welding current/A Arc length/mm
Fig. 4 — Demonstration of experimental setup. The sensing system for the experiment is identical with that in Fig. 1. A computer connected to the camera is used for image processing, weld pool reconstruction, characterization, and to calculate the current output using the human welder response model.
54 2, 3, 4, 5
To define these parameters, the 2D parametric model of the weld pool shown in Fig. 3A x = ±ay b (1 − y ) , (a > 0, 1 ≥ b > 0) r
r
r
(1)
is adopted (Ref. 28). This model uses x r = x/L and y r = y/L. Once this model is obtained, the width of the weld pool is then calculated b b b w = w × L = 2 aL (2) r 1 + b 1 + b
Figure 3B shows the longitudinal intercepted area of the weld pool in oxy plane. The convexity is defined as the intercepted area divided by the length of the weld pool. Modeling the human welder response is then to correlate his adjustment ∆ I k as a function of the characteristic parameters in different instants around instant k. This can be done using the standard least squares algorithm. To obtain this optimal model, F-test (Ref. 31) has also been used to determine the instant range that needs to be included in the model for each of the characteristic parameters. As a result, the following model was obtained (Refs. 18, 19):
∆ I
k
− 0.4725 ∆I + 0.1366∆I k −1 k −2
= 0.6097 L
− 2.2283 L k
+ 1.6137 L
− 1.2675 W
k −3 k −5
+0.0930 W k
−5
C + 30.3658C
−4
k −3
+ 1.7667 W
− 0.6088 W k
k −3
−6
+ 19.6357C
k −4
−67.6373C k − + 18.7761C k − 5
k −4
6
(3)
where ∆ I k–j is the current adjustment at instant k–j with a 0.5-s sampling period. It can be found the human welder adjusts the current based on the previous current adjustments and weld pool surfaces. That
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A
A
B
B
C
C
Fig. 5 — Results from experiment with initial current of 50 A. A — The cur rent and voltage; B — the backside weld bead width; C — the backside backside weld bead (the unit of x and y axis is pixel).
Fig. 6 — Results Results from experiment with initial current of 54 A. A — The cur rent and voltage; B — the backside weld bead width; width; C — the backside weld bead (the unit of x and y axis is pixel).
is, the adjustment on the welding current by the human welder requires the length, width, and convexity of the weld pool surface to model adequately. In addition, the human welder makes the adjustment on the welding current based also on the pre vious adjustments he made 1 s ago.
Experimental System and Methods
Fig. 7 — Diagram of control system of the human welder response model with additional additional low-pass filter. filter.
In this section, the experimental setup and methods used to implement the human welder response model-based control are summarized. Experimental Setup
The configuration of the experimen158-s MAY 2013, VOL. 92
B
A
C
D
Fig. 8 — Curren Currentt and voltage of the the experiments experiments.. A — Initi Initial al current current of 50 A; B — initial initial curren currentt of 54 54 A; C — initial initial current current of 58 A; D — init initial ial current current of of 62 A.
tal system is shown in Fig. 4. As mentioned in the human welder response model section on a vision-based sensing subsystem, having the laser pattern pro jecting the dot matrix on the weld pool surface, part of the dot matrix pattern is specularly reflected from it. Intercepted by the imaging plane, the reflection pattern is then captured by the camera. A computer connected to the camera is responsible for processing the captured image, reconstructing the weld pool surface, and extracting the characteristic parameters. Based on the obtained characteristic parameters of the weld pool surface, the adjustment needed for the weldin wel dingg cur curren rentt is cal calcul culate ated d by the human welder response model. Accordi Acc ording ng to to the the princ principle iple of weld welders’ ers’ behavior briefed previously, a welder starts a welding weld ing perfor performanc mancee with an optima optimall estiestimation of the welding parameters based on past experience. To imitate the welder’s be-
Table 5 — Welding Parameters for Root Opening Robustness Experiments
Welding parameters Root opening/mm Arc length/mm length/mm Initial current/A
1 0 54
havior, in each experiment of the study, specific welding conditions (welding conditions and parameters that are not changed/ad justed just ed on on purpos purposee in each part particul icular ar exper exper-iment including welding speed, arc length, etc.) and an initial current are first applied for the weld pool to grow freely to complete joint pen penetra etration tion.. Then the weld welding ing proc process ess is manually switched to control mode, i.e., the human welder response model starts to adjust the current for consistent complete joint pen penetra etration tion..
Experiments 2 2 3 58
3 [0, 5] 54
Experimental Approach
In a manual welding process, a qualified welder can control the welding process to obtain a nearly uniform penetration (backside weld bead width) that he/she desires, even with different welding conditions. To this end, a number of experiments were conducted with different welding conditions and initial welding current in this study. At the beginning of o f each experiment, the weld pool grows freely.
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H C R A E S E R G N I D L E W
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W E D L Fig. 9 — The backside appearance of the weld bead. A — Initial current of 50 Fig. 10 — The backside width of weld beads with a different initial current. B — initial current of 54 A; C — initial current of 58 A; D — initial initial current D A; of 62 A. I N G R With specific welding conditions and ini- mates an initial welding current to start a rent quickly and accurately to maintain current, the welding process is able to manual welding operation. The past expeuniform penetration, or produce an unE tial to complete joint penetration. Yet, rience-based estimation might vary within qualified weld with a same failure pattern. S reach the dimension of the weld pool at coma reasonable range. Also, the arc length to verify the effectiveness of E plete joint penetration in each experiment maintained by the welder might not always Therefore, the model is to check if it is able to control A is expected to be different. Then the ex- be the same during manual welding, as the welding process to a comparatively opening ning . The weld ing consistent penetration, i.e., the backside R periment is manually switched to control well as the root ope that is, to apply the human welder speed, on the other hand, does not change width of the weld bead under differ different ent C mode, response model to control the process. much when controlled by the welder, al welding conditions and initial current can H Specifically, the model adjusts the welding though it might vary within a small range. converge to a constant within a small varicurrent based on the geometry of the 3D weld pool surface such that the adjusted welding current controls controls the process to obtain a desired penetration that is evaluated by the backside weld bead width. After each experiment, the width of the obtained backside weld bead is measured to veri fy the effe ctive nes nesss of the hum human an welder response model-based control. The experimental parameters used here are listed in Table Table 1. The pipe used in this study is 4-in. nom. stainless T304/304L Schedule 5. The initial current is in 50, 62 A, the arc length varies within 2, 5 mm, and the joint opening changes from 0 to 5 mm. The rotation speed of the pipe, i.e., the welding speed and up-down motion of the torch are controlled by the computer to achieve the required welding speed and arc length. The effectiveness and robustness of the human welder response model-based control will be evaluated against those welding parameter variations in this paper. As ment mentioned ioned befor e, a welde welderr esti160-s
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The welding speed is constant for the experiments in this study. The welder’s behavior under large welding speed variation is the authors’ future work and beyond the scope of the first study of this kind. As presented in the introduction, the response model is developed based on the behavior of a human welder with limited skills. Given the physical limitation as a human, the welder might feel stress, fatigue, and lack of concentration in manual welding. Because of the possible inconsistent welding behavior, the welder’s response data used to develop the model cannot represent the prime performance of the welder. In this sense, the model is only able to present an average performance of the human welder. However, the control based on the human response model should be able to get rid of the inconsistency, which is a major issue with manual welding. The model is expected to either consistently produce a good weld by adjusting the cur-
ation margin. There are several types of relevant data in this study. For the process, such data include welding current, arc length, and welding speed. Since the human welder response model under this specific study controls the welding process only by ad justing the welding current, such data are especially concerned with analyzing the performance of the control. Second, for the weld pool surface, the data include all its characteristic parameters, i.e., the length, width, and convexity. Studying the current adjustment and variation in the characteristic parameters can reveal how the model controls the welding process. At last for the weld bead, the backside width is the major data of interest. All the these se dat data, a, exce pt the back side weld bead width, are acquired/recorded in real time. The backside bead width is measured with one sample/s interval offline. For example, if the welding speed in one experiment is 1 mm/s, then the bead width is measured every 1 mm while it is
A
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H C R A E S E R G N I D C D L Fig. 11 — Current and voltage from arc length robustness experiments. A — Arc length of 2 mm; B — arc length of 3 mm; C — arc length of 4 mm; D — arc E length of 5 mm. W measured every 1.5 mm if the welding speed is 1.5 mm/s, etc. In this sense, the bead width measurement can be matched with the time scale for other types of data.
Human Welder Response Model Control In this section, the results from the experiments with initial current are presented and analyzed. As the first time to implement the human welder response model-based control, we focused on how the model controls the process to consistent complete joint penetration. Major welding parameters used in the experiments are listed in Table 2; the rest are the same as in Table 1. Results from the two experiments, including the current, backside bead appearance, and its width, are presented in Figs. 5 and 6, respectively.
The data acquisition starts before the welding process begins. It is found that at the beginning of the two experiments, there are about 13-s (shown in Fig. 5A) and 12-s (shown in Fig. 6A) periods during which the current is 0, and the voltage is the open-circuit one at approximately 70 V. The red vertical dash line in Fig. 5A indicates the time instant when the process is switched to the model-based control mode, i.e., the human welder response model is applied to adjust the current based on the 3D weld pool surface. From the figure, the control begins at 40 s and ends at approximately 96 s. The length of the weld bead obtained in this period can be easily calculated using the known welding speed. Since the position of the end of the welding is clearly seen in Fig. 5C, the position where the welding process begins to be controlled on the produced weld can
be determined as shown by the red vertical line in the middle of Fig. 5C. By the start position of the model control, each of the weld beads in the two experiments is divided into two zones as shown in Figs. 5C and 6C. In zone A, the backside bead width is determined by the welding conditions used and initial current. In the first experiment, the initial current is 50 A. The average width of the backside bead in this zone, shown in Fig. 5B, is about 1.7 mm. With a greater initial current (54 A) in the second experiment, the average width becomes 3.2 mm in zone A as can be seen in Fig. 6B. In zone B, the human welder response model starts to control the process for a desired and consistent penetration. Despite the fluctuation, the average width for the first experiment, shown in Fig. 5B as about 4.8 mm, and that in the second experiment, shown in Fig. 6B as about 4.7 mm, are considered WELDING JOURNAL 161-s
tem and methods section, the model represents the average performance of a welder with limited skill. As a result, it shows in the human welder response model control section that the model is not able to adjust the current accurately enough to reduce the oscillation although a comparatively consistent penetration is obtained. It is known that a skilled welder can avoid the current ripple with smooth current adjustment. In this sense, to smooth the model’s adjustment is to filter out the high-frequency part of the calculated current adjustment, i.e., to neutralize the dynamics associated with the overreaction in the response model. A simple method is to adapt a digital low-pass filter after the model in the control system as shown in Fig. 7. The low-pass filter used in this study can be written in Equation 4.
A
B
C
∆ I ′ k
W E L D D I N G R E Fig. 12 — The backside weld beads from arc length robustness exp eriments. A — Arc length of 2 mm; S B — arc length of 3 mm; C — arc length of 4 mm; D — arc length of 5 mm. E A the same. the fluctuation by smoothing the beR It can be found there are current fluc- avoid havior on the current adjustment. C tuations during the control period in both Although there are similar current H experiments, shown in Figs. 5A and 6A. fluctuations shown in both experiments, it That means the model is able to adjust the current quickly but not skilled enough to reduce the current ripples. That leads to noticeable oscillations of the backside bead width, which are clearly shown in Figs. 5B, C and 6B, C. The current fluctuates between 64 and 55 A in the first experiment, and 64 and 54 A in the second. Correspondingly, the backside bead width changes from about 5.8 to 3.5 mm in the first experiment, and 5.2 to 3.7 mm in the second experiment. The fluctuation of the current adjusted by the model is understandable since the model is developed using the data from the behavior of a novice welder with limited skills. The reason for using an unskilled welder is that the authors intend to study and follow the development of welder skills and responses. It is a common welding scenario that an unskilled welder cannot predict the process quickly and accurately so that he/she would frequently overreact or underreact to the welding process. A seasoned welder can easily
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can be found that the current, controlled by the model, is settled down within a certain amount of time. This is because the human welder response model is stable since all the poles of the model (Equation 3) are inside the unit circle in the Z plane (Ref. 29). That means the model can control the process to a steady state. This is understandable that any welder with limited skill should be able to produce a steady welding process. Difference in the width of the backside bead in the two experiments in zone A indicates the different characteristic parameters of the 3D weld pool surface are obtained. Despite the difference, the welding processes in the two experiments reach a nearly identical backside bead width (4.8 and 4.7 mm) after the control of the human welder response model.
Improvement of the Human Welder Response Model As discussed in the experimental sys-
= α∆ I k–1 + (1 – α )∆ I ′ k, 0<α<1 (4)
where ∆ I ′ k and ∆ I ′ k–1 are the filtered current adjustment at time instant k and k–1, respectively, and ∆ I k is the current adjustment calculated by the welder response model at time instant k. Coefficient α controls the frequency bandwidth of the filter. A greater α gives a wider bandwidth. In this study, α is selected to be 0.5. Since the filter blocks high-frequency components in the current output, its function would be pronounced during the transition period. However, when the current approaches its steady state, the highfrequency components become insignificant. The steady-state value of the current for a particular experiment is not affected by the filter. Moreover, since the current adjustment is smoothed by the filter, the current ripple is expected to be minimized. The backside weld bead width is expected to be more consistent. The transiting period also should be reduced significantly. In this sense, adapting a lowpass filter to the human welder response model makes the model function like a more skilled welder.
Results and Analysis of Robustness Experiments The human weld response modelbased control is now improved simply by adding a low-pass filter as schematically illustrated in Fig. 7. To confirm its effectiveness in controlling the process to achieve the desired weld penetration, various experiments were designed and conducted in this section using this improved system to examine its performance/robustness under different welding conditions and initial welding currents. In the first subsection, the experiments with different initial current amperages are conducted. The robustness of
A
B
Fig. 13 — The backside width of weld beads from arc length robustness experiments.
the human welder response model-based control with respect to the initial current is analyzed. In the next subsection, the arc length changes from 2 to 5 mm in the conducted experiments. The root opening is designed to vary from 0 to 5 mm in the last subsection. Robustness with Respect to Initial Current
Experiments with different initial welding currents are conducted again but with the improved control system. Since the purpose is to examine the effectiveness of the improved control, more initial currents (Table 3) in a greater range are used to examine the system’s robustness against the initial current used. The results are shown in Figs. 8 to 10. Figure 8 shows the current and voltage from these four experiments; the backside weld beads obtained are demonstrated in Fig. 9; and the measurements of the backside width of the weld beads are presented in Fig. 10. Since the initial welding current determines the backside width of weld bead in zone A, with different initial currents, the obtained weld beads have different backside widths. Specifically, the backside bead width in the experiment with the initial current 50 A is 3.2 mm; with greater initial current (54 A) in the second experiment, the backside width is 3.7 mm; the backside width obtained in the third experiment (initial current 58 A) and fourth (initial current 62 A) are 4.7 and 5 mm. The settling time (set the error margin 5%) (Ref. 30) for the process to achieve the steady state is in the range from 3.5 to 4.5 s. For the experiment with an initial current of 62 A, the steady-state current is only about 0.5 A less than the initial current. The transition period is negligible from Fig. 8D. Besides that, the settling times of the other ex-
C
periments are close to each other, only with 1 s difference, despite the difference in the initial current used. As discussed in the D human welder response model section under the principle of Fig. 14 — Front side of the weld joint before experiment for root opening ro bustness experiments. A — 0 nominal root opening; B — 2-mm nominal human welder’s be- root opening; C — nominal root opening increases fro m 0 to 5 mm (0.21 havior, a welder would in.); D — close review of the wide opening for marked area in C. start to weld with an optimally estimated initial current (initial welding parameters) tween the welding current and backside based on past experiences. A skilled welder bead width is understandable. There are might adjust the initial current close to the two dynamic processes involved, and each steady-state current that produces a desired of them has a settling time. First, after the penetration. The effect from the transition current settles down reaching its steady period on the result then can be minimized. state, the weld pool surface will take an ad An unskilled welder, on the other hand, ditional transition period to reach the cormight not be able to predict the initial curresponding dimension. The first dynamic rent close to a current that leads to the deprocess is the transfer from the welding sired penetration. current (parameters) to the weld pool surThe transition period for the backside face. Second, after the 3D weld pool sur weld bead is different such that the settling face (front side) is settled down and time for the backside bead width differs reaches its steady state, the 3D weld pool from that of the welding current. From surface on the backside that determines Fig. 8C and D, the initial current is close the backside bead width will take extra enough to the steady-state current. Theretime to reach its corresponding steady fore, the width of the backside bead in state. The transition period associated each of the two experiments only increases with the backside width of weld bead is about 0.2 mm after the control is applied longer than the transition time of the as shown in Fig. 10. The effect from the welding current. settling time on the welds produced for Figure 8 shows the steady-state current these two experiments is negligible. Howcorresponding to the penetration obever, the settling time for the backside tained by the human welder response bead width observed from the experimodel varies from 59.5 to 61.5 A. And the ments with an initial current of 50 and 54 steady-state width of the obtained weld A is about 13 and 10 s, respectively. The beads, shown in Fig. 10, converges within backside weld beads take a longer time the range from 5.0 to 5.3 mm in zone B. A than the welding current to reach the 0.3-mm deviation in the backside bead steady state. width is considered acceptable. It is known The difference in the settling time bethat the weld pool is dynamic and vibratWELDING JOURNAL 163-s
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ing through the welding process. Even a constant welding input might cause the backside bead width varies within a small range. It is possible that the difference of the weld pool dimension generated by the current varying from 59.5 to 61.5 A is unperceivable to the human welder response model. In this sense, despite the different backside width obtained at the beginning of those experiments, because of different initial current, a consistent penetration with only a 0.3-mm width variation is achieved using the model-based control. Robustness with Respect to Arc Length
Arc length is another variation whose effect on the control system needs to be examined. Hence, experiments with different arc lengths are conducted. The initial current, arc length, and welding speed are listed in Table 4; the rest are the same as in Table 1. As can be seen from Table 4, 164-s MAY 2013, VOL. 92
B
Fig. 15 — Current and voltage from joint opening robustness experiments. A — 0 nominal joint o pening; B — 2-mm nominal joint opening; C — nomi nal joint opening increases from 0 to 5 mm (0.21 in.).
only the arc length differs in the four experiments to be conducted while other parameters are the same. The resultant current, backside weld bead, and backside width measurements are presented in Figs. 11–13, respectively. With an increase in the arc length, the arc distribution becomes broader, and the arc energy intensity decreases. The penetration capability thus reduces. From Figs. 12 and 13, the width of the four weld beads in zone A reduces down from 3.7 to 2.5 mm. To obtain a consistent penetration, the model-based control increases the steady-state current from 61 to 64 A in the experiments with the arc length changing from 2 to 5 mm. As mentioned in the results and analysis of robustness experiments section under robustness with respect to initial current, a 2A deviation of the steady-state current is considered a reasonable margin for the control of the human welder response mode. Only a 3A deviation is obtained here because of the difference in arc lengths. Therefore, the arc length does not significantly affect the value of steady-state current during the control using the human welder response model. It is noticed that the voltage in Fig. 13 increases from 9 to 9.8 V due to the increase in the arc length. A difference of 0.8
V in the arc voltage is observed between 5 and 2 mm arc length. As for the settling time for the current in the four experiments, it is similar to each other, which is about 5 s, as can be seen from Fig. 11. The transition period for the backside width is also similar to each other, as shown in Fig. 13. The arc length difference does not affect that transition period of the current or backside width of the weld bead. From Fig. 13, the backside width of the weld bead in the four experiments converges to about 5.2 mm in zone B. Among the four weld beads, the one with 2 mm arc length reaches the largest steady-state width (5.4 mm), and the weld beads with 3 and 4 mm arc length have the smallest steady-state width, which is about 5.0 mm. From Figs. 13 and 10, one can find that a nearly identical backside width, which is about 5.2 mm with a 0.4mm variation margin is obtained despite a difference in arc length and initial current. That means the human welder response model is able to control the welding process to achieve a consistent penetration under different arc length and initial current. It is reasonable because the model changes the current based on its previous current adjustments and the 3D weld pool geometry, as discussed in the human welder response model heading under the human welder response model section. Differences in weld pool dimension caused by different initial currents or arc lengths might not be undetectable for the human welder response model. Therefore, the model is able to maintain a consistent penetration despite the different arc length and initial current.
Root Opening Robustness
Root opening is difficult to be precisely controlled in production. The effectiveness of the human welder response modelbased control needs to be examined under varying/different openings. In this subsection, experiments with different root opening conditions/variations are conducted. The root opening, arc length, and welding speed are listed in Table 5; the rest are the same as in Table 1. There are three experiments. The nominal/intentional root opening in the first experiment is 0 mm as shown in Fig. 14A. As demonstrated in Fig. 14B, a 2-mm root opening is used in the second experiment. In the third experiment, the nominal opening gradually increases from 0 to about 5 mm as shown in Fig. 14C and D. The resultant current, backside weld bead, and their widths are presented in Figs. 15–17, respectively. The steady-state current differs in these three experiments. The root opening is close to zero in the first experiment. The welding process is close to those in the results and analysis of robustness experiments section under robustness with respect to initial current and robustness with respect to arc length. Therefore, the steady-state current (63 A) is close to the resultant steady-state current obtained in the last two subsections. However, as the root opening increases, the weld pool surface tends to be more concave, which means the convexity of the weld pool is smaller. The current adjustment controlled by the human welder response model (Equation 3) tends to be smaller accordingly. The penetration capability of the arc also increases with the opening. Therefore, less heat input is required to produce the same penetration as the opening increases. The steady-state current for the second and third experiments are reduced to 61.5 and 54 A, respectively. The obtained backside widths of the weld beads in these three experiments are about 5.2, 5.5, and 5.6 mm, which are considered consistent with a reasonably small variation margin. It needs to be mentioned that the welding process stops at the position where the root opening is about 4 mm in the third experiment. The experimental results in experiment 3 only claim that the human welder response model-based control can control the welding process to maintain the consistent penetration for a root opening 4 mm or smaller.
Conclusion This paper addresses implementing the human welder response model to ad just the welding current in reply to the characteristic parameters of the 3D weld
A
B
C Fig. 16 — The backside appearance of the weld beads from root opening robustness experiments. A — 0 nominal root opening; B — 2-mm nominal root opening; C — nominal root opening increases from 0 to 5 mm (0.21 in.).
pool surface for maintaining consistent, complete joint penetration in GTAW. The effectiveness and robustness of the human welder response model-based control are verified in the experiments with different welding conditions and the initial current. The material used in the experiments is stainless steel pipe (4-in. nom. stainless T-304/304L Schedule 5). For the initial conditions, the current varies from 50 to 62 A, arc length is within [2, 5 mm], and the root opening changes from 0 to 5 mm. Under the experimental conditions used, the following were found: • The human welder response model can diminish the inconsistent manual welding performance. The model can control the welding process by adjusting the current to maintain consistent complete joint penetration. • The backside width of the weld bead corresponding to the consistent complete joint penetration is about 5.2 mm with a 0.4-mm error margin. • The maximum root opening at which the human welder response model can produce a consistent complete joint penetration is about 4 mm. The effectiveness and robustness of the human welder response model control are verified against different welding conditions and initial current amperages. Future studies will focus on modeling behaviors of skilled welders. The resultant models may be directly used without lowpass filters to develop control systems for improved performance. Differences and similarities with those of the novice welder
H C will be analyzed and used to help train and R A improve less skilled welders. E Acknowledgments S E This work is funded by the National Science Foundation under grant CMMI- R 0927707. We wish to thank Yi Lu and G Yukang Liu for their assistance on experiments and graphics, and Lee Kvidahl for his N I technical guidance on manual pipe welding. D L References E 1. Uttrachi, G. D. 2007.Welder shortage requires new thinking. Welding Journal 86(1): 6. W 2. Welding shortage fact sheet. American Welding Society. 2007. http: //file s.aws.org /pr/ shortagefactsheet.pdf . 3. Richardson, R. W., and Edwards, F. S. 1995. Controlling GT arc length from arc light emissions. Trends in Welding, Proceedings of the 4th International Conference, pp. 715–720. 4. Li, P. J., and Zhang, Y. M. 2011. Precision sensing of arc length in GTAW based on arc light spectrum. Journal of Manufacturing and Science Engineering 123: 62–65. 5. Stecker, S. D. 1996. Characterization and application of weld pool oscillation phenomenon for penetration control of gas tungsten arc welding. MS thesis. The Ohio State University, Department of Welding Engineering. 6. Hartman, D. A., DeLapp, D. R., Barnett, R. J., and Cook, G. E. 1998. A neural net work/fuzzy logic system for weld penetration control. Trends in Welding Research, Proceedings of the 5th International Conference, pp. 1096– 1101. 7. Balfour, C., Smith, J. S., and AlShamma a, A. I. 2006. A novel edge feature correlation algorithm for real-time computer vision-based molten weld pool measurements. ′
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Fig. 17 — The backside width of weld beads from root opening robustness experiments.
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