laser welding experiment

Plastic 3D Printing Service

Fused Deposition Modeling

HP Multi Jet Fusion

Selective Laser Sintering

Stereolithography

Production Photopolymers

Nexa3D LSPc

Direct Metal Laser Sintering

Metal Binder Jetting

Vapor Smoothing 3D Prints

CNC Machining

CNC Milling

CNC Turning

Wire EDM Machining

Medical CNC

CNC Routing

Sheet Metal Fabrication

Sheet Cutting

Laser Cutting

Waterjet Cutting

Plasma Cutting

Tube Bending

Laser Tube Cutting

Laser Engraving

Custom Die Cutting

Plastic Injection Molding

Quick-Turn Molding

Prototype Molding

Bridge Molding

Production Molding

Overmolding

Insert Molding

Compression Molding

Urethane and Silicone Casting

Plastic Extrusion

Injection Molded Surface Finishes

Custom Plastic Fabrication

Micro Molding

Die Casting

Metal Stamping

Metal Extrusion

Metal Injection Molding

Assembly Services

Rapid Prototyping

High-Volume Production

Precision Grinding

Surface Grinding

Powder Coating

Aerospace and Defense

Consumer Products

Design Agencies

Electronics and Semiconductors

Hardware Startups

Medical and Dental

Supply Chain and Purchasing

All Technical Guides

Design Guides

eBooks Library

3D Printing Articles

Injection Molding Articles

Machining Articles

Sheet Cutting Articles

Xometry Production Guide

CAD Add-ins

Manufacturing Standards

Standard Sheet Thicknesses

Standard Tube Sizes

Standard Threads

Standard Inserts

Case Studies

ITAR and Certifications

Supplier Community

Release Notes

Call: +1-800-983-1959

Email: [email protected]

Discover Xometry Teamspace

Meet An Account Rep

eProcurement Integrations

Bulk Upload for Production Quotes

Onboard Xometry As Your Vendor

How to Use the Xometry Instant Quoting Engine®

Test Drive Xometry

Tool Library and Tool Detail Pages

Part Revisions & Same-Suppliers for Repeat Orders

Xometry's Privacy and Security

Xometry’s Quality Assurance

Xometry’s Supplier Network

Xometry's Machine Learning

Xometry's Manufacturing Community

Laser Welding 101

.css-2xf3ee{font-size:0.6em;margin-left:-2em;position:absolute;color:#22445F;} .css-14nvrlq{display:inline-block;line-height:1;height:1em;background-color:currentColor;-webkit-mask:url(https://assets.xometry.com/fontawesome-pro/v6/svgs/light/link.svg) no-repeat center/contain content-box;mask:url(https://assets.xometry.com/fontawesome-pro/v6/svgs/light/link.svg) no-repeat center/contain content-box;-webkit-mask:url(https://assets.xometry.com/fontawesome-pro/v6/svgs/light/link.svg) no-repeat center/contain content-box;aspect-ratio:640/512;vertical-align:-15%;}.css-14nvrlq:before{content:"";} What Is Laser Welding?

Laser welding is a precise and surprisingly delicate welding process that uses beam light to join metal parts. The laser beam is produced by a freestanding laser source and directed to a machine-mounted or hand-held collimator ‘torch’ and onto the workpiece. The directed energy beam heats and melts the workpieces (and filler rod if required), and the resulting melt pool fuses the targets together, forming a well-integrated weld.

A key advantage of laser welding is the high precision and close control of applied energy. The energy can be precisely directed to the exact location where the weld is needed with virtually no over-application or spread. It allows for extremely precise liquefaction with a much smaller heat-affected zone than any other welding method. This causes less damage to surrounding areas and reduces bulk heating and the associated deformation.

Laser welding is an increasingly important technique in high-tech manufacturing and engineering. It offers many advantages over traditional welding techniques. Laser welding can create clean, strong welds quickly and precisely. It is commonly used in the automotive, aerospace, and medical industries, as well as in electronics manufacturing. It is particularly useful for welding materials like aluminum that are difficult to weld by traditional methods.

The process was first developed in the ‘60s, almost in parallel with the development of lasers themselves. The first experiments in laser welding were conducted by researchers at the Bell Telephone Laboratories in the US. The first laser welds were made using a ruby laser, applying short pulses of high-intensity energy in coherent beams. When pulses were focused onto a small spot at the junction of metal parts, the metals melted and flowed together. Even in these early experiments, this resulted in very narrow and precise welds, with minimal HAZ (heat-affected zones) and distortion. It even demonstrated some ability to join dissimilar materials.

How Laser Welding Works

Laser welding is a technique that uses a focused collimated high-intensity beam of light to melt and fuse metal parts, sometimes with extra material from a filler rod. The process works by generating laser light and then delivering it to a collimator/optics head. It is then focused onto the junction of the metal parts, causing a highly localized heat buildup and restricted melt pool.

The laser welding beam is typically generated by a solid-state, fiber, or CO2 laser, each of which has relative advantages. At the beam’s focus point, the metal reaches its melting point and forms a localized pool, into which the filler rod can be melted as required. The laser beam is then moved along the surface of the joint. This melts a leading edge and leaves the molten, fused trailing edge of the pool to cool and solidify. In a successful weld, the cooled metal attaches to both parts to a roughly equal degree and remains free of oxidation. 

The Basics of the Laser Welding Process

These are the generic steps in the laser welding process:

• Clean the parts to be welded and position them accurately. The contact line should be closed and gap-free to improve weld quality.
• Use manual clamps or automated fixtures to hold parts in place and keep them stable during the welding process.
• Adjust the beam’s focal point onto the welding area. The optical gear in the welding torch usually provides for easy adjustment of focus.
• Adjust the beam power and test it on scrap material and trial parts. Before moving to the workpiece, make sure it is putting out sufficient energy to melt the material but not enough to excessively heat parts. 
• Apply the beam at the start of the welding area. Once an appropriate melt pool has formed, it must be traversed along the weld in a steady motion. Traditional welding techniques such as hot point rotation will encourage good fusion and improve the weld quality.
• Cool the part naturally once the welding is complete. You can also quench it in water or use other cooling methods.

Kinds of Materials That Can be Laser Welded

The most common materials that can be laser welded are listed below:

• Metals: Examples are: aluminum, copper, brass, steel, titanium, and nickel. The process can be used to join pieces of significantly divergent thicknesses, increasing its applications to a wider selection of tasks than traditional thermal or electrical welding methods.
• Plastics. Lasers can be used to weld some thermoplastics, including: polycarbonate, nylon, and ABS. Low heating and highly localized melting result in quality welds.
• Ceramics. Some ceramics (particularly alumina and zirconia) can be laser welded. These and some other ceramics can be melted and fused via laser in a way that is much harder to achieve by normal thermal means.
• Composites. Carbon fiber-reinforced plastics (CFRPs) are amenable to this technique. There is also advanced research and early success in laser welding metal parts to carbon fiber composites.

The suitability of a material for laser welding depends on its physical properties, such as melting temperature, albedo, thermal conductivity, and its propensity to melt without charring. Significant experience and careful laser frequency selection may be needed for highly reflective materials.

The Challenges With Laser Welding

These are the common challenges seen in laser welding:

• Material Selection: Some materials, such as highly reflective metals, are difficult to weld with lasers because light naturally reflects away from their surfaces. Similarly, some plastics and composites are also difficult to weld due to their low thermal conductivity.
• Joint Preparation: Proper joint preparation is critical for a successful laser weld. The joint surfaces must be free from contaminants and properly aligned. Any misalignment or gaps in the joint can result in incomplete welds or weak joints.
• Process Control: Laser welding is a highly automated process and maintaining tight process control is essential for producing consistent, high-quality welds. The laser power, speed, and focus must be carefully controlled to achieve the desired weld characteristics.
• Safety: Laser welding can pose safety risks if not properly controlled. The intense light and heat generated by the laser can cause eye and skin damage. Safety measures such as proper eye protection must be in place to protect the operator and any nearby personnel.
• Cost: The initial cost of equipment for laser welding can be high, making it less accessible to smaller businesses or operations. Additionally, maintenance and repair costs can also be significant, which can add to the overall cost of laser welding.

Different Types of Laser Welding

These are the most common methods of laser welding:

• Conduction Welding: This method uses the lowest power rating of any laser-based approach. It merges the melted edges by capillary action alone, with no filler. This approach is best suited to welding precisely fitted edges of thin materials.
• Deep Penetration Welding: This method is suitable for welding thicker materials. It uses high laser power to heat a deep and wide portion of the material. In general, the laser is first used to cut a keyhole that penetrates through the material (ensuring full-thickness welding). The resulting hole is then closed with a molten filler rod at its trailing edge, as the laser progresses along the weld.
• Laser Spot Welding: This method is best used for small, complex parts. The laser creates small, localized welds. These spot welds can make point joints between edges, or melt through one part to merge with the part below.
• Laser Seam Welding: This approach makes long, continuous seams. It often uses a filler rod to create a filet at the joint using similar pool control motions to those seen in electrical and traditional thermal methods.
• Hybrid Laser Welding: This method mixes laser and other welding processes such as MIG and TIG. Combining processes in this way can give you the advantages of both systems. 

Types of Lasers Used in Laser Welding

The types of lasers used in laser welding are listed below:

1. CO2 (Carbon Dioxide) Lasers

Carbon dioxide lasers are mainstays in welding equipment thanks to their high power output and small spot size when focused. They operate in the mid-infrared emission range and are capable of welding most materials, although the reflection of initial power can make the onset of melting slowly in stainless steel, titanium, and some other reflective metals.

2. Nd:YAG (Neodymium-doped Yttrium Aluminum Garnet) Lasers

Nd:YAG (neodymium-doped yttrium aluminum garnet) solid-state lasers are also commonly used for welding. This laser type generates high-powered infrared light with a wavelength of 1.064 micrometers. It is a good option because metallic materials absorb this wavelength better than others in the infrared spectrum. Nd:YAG setups are thus particularly useful in welding aluminum, stainless steel, and titanium alloys. Because it combines high energy output and good focusability with minimal maintenance requirements, this system is commonly used in industrial applications, such as automotive and aerospace manufacturing. 

3. Fiber Lasers

Fiber lasers are good options, delivering high power, superior beam quality, and electrical power efficiency. The laser energy comes from a laser diode. It is transmitted through a fiber optic connection to a collimating/focusing torch that can be easily directed to the welding site.

This laser welder type integrates well into automated equipment. It has a long device life expectancy and low maintenance needs. For more information, see our how do fiber lasers work guide.

4. Disk Lasers

Disk lasers are alternative forms of solid-state lasers that are beginning to be used for welding. In a disk laser, the laser medium is a thin, liquid-cooled disk of laser-excitable semiconductive material that is pumped by several laser diodes. The output beam can be transmitted in a rigid reflector path or through a light pipe to the collimator/focus torch. They’re valuable because of their high power capacities, good beam quality, efficient cooling, low maintenance requirements, and long functional life expectancy.

Key Parameters That Affect the Quality of a Laser Weld

• Laser Power: This is the amount of energy delivered by the laser to the workpiece. Higher laser power allows faster welding and greater penetration, but good control of power levels is critical in achieving good welds.
• Spot Size: A higher-quality beam, better collimation, and better-quality focus optics will result in a smaller laser point at the weld. This delivers higher effective power and more controlled melt/weld progression.
• Feed/Traverse Speed: The rate (and motion pattern) of the laser as it passes along the seam defines the weld quality, the size of the HAZ, and the level of distortion in parts.
• Shielding Gas: Inert gasses are used to prevent weld oxidation. The type and flow rate of shield gas must be suitable for the weld and materials.
• Material Thickness: This is a critical parameter in that there are limitations to what any particular laser welder can deliver. As part thicknesses increase, laser power must rise and feed rates usually drop. Eventually, you reach the limits of the welder’s capabilities. 
• Joint Design: Joint design affects weld quality significantly. Close conformance/fit between parts and accessible weld positions are very beneficial in weld quality.

What Are the Main Applications of Laser Welding?

Laser welding is used in a wide range of industries:

• Automotive. Body panels, engine components, suspension parts, fuel injectors, and sensors can all be joined or constructed using laser welders.
• Aerospace. Laser precision is helpful in aircraft engines, landing gear, and other components. It is also valuable in the manufacture of rockets and spacecraft.
• Medical. Medical devices such as pacemakers, dental implants, surgical instruments, surgical implants, and prosthetics all need detailed welds.
• Electronics. Small, precise welders are important to the manufacture of circuit boards, complex component packages, sensors, smartphones, laptops, and more.
• Jewelry. Both manual and automated laser welding is used in the manufacture of jewelry to create intricate designs.
• Mold Tools and Dies: The tools and dies used in manufacturing processes can be difficult to repair. The minimal excess heat created by a laser weld means there’s less finishing work to be done after a repair. 

The Advantages and Disadvantages of Laser Welding

The advantages of laser welding are listed below:

• Precision: Laser welding is a precise welding technique that can create small, intricate welds with high accuracy. The beam energy can be controlled very precisely, minimizing the heat-affected zone and keeping distortion and material waste to a minimum. 
• Speed: The technique is fast. Since the energy is very concentrated, it heats a melt pool quickly. Heat doesn’t have time to spread as far as it does under other welding methods. The pool’s leading edge can be advanced quickly using otherwise standard welding practices. 
• Versatility: Laser welding can be used to join a wide range of materials, including metals, plastics, and even some ceramics. The process can also be used to weld dissimilar materials together, which is generally impossible with other welding techniques.
• Quality: Laser welding results in high-quality welds with consistent mechanical properties. If no filler rod is required, welds will generally match the properties of the joined materials in terms of strength, durability, and corrosion resistance. The welds tend to be free from defects such as porosity, inclusions, and fractures.
• Automation: Laser welding is easy to automate, delivering high-volume production and consistent quality in ways that previously could only be approached by spot welding. This is particularly important in the automotive, aerospace, and electronics industries, where precision and repeatability are key.

The key limitations of laser welding are:

• Equipment Cost: Laser welding equipment costs more than that for electrical or traditional thermal welding.
• Safety: Lasers can be hazardous, so they require careful management and good safety practices to prevent burns or eye injuries.
• Material Limitations: Though laser welding is effective on a vast range of materials, many plastics and ceramics cannot be welded for thermal and chemical reasons.
• Weld Geometry: Laser welding is best suited to thin, light, and precise applications with very close-fitting parts that the optical parts can easily access. It is not currently well adapted to heavy-duty roles such as shipbuilding.
• Joint Preparation: The process is less forgiving of contaminants, surface oxidation, and gaps between parts than traditional processes.
• Edge Preparation: Material edges must be close-fitting and smooth for good results.
• Maintenance: Some types of laser welding equipment require intensive maintenance and setup, increasing operational costs.

This article presented laser welding, explained it, and discussed the process in detail and its various types.

The content appearing on this webpage is for informational purposes only. Xometry makes no representation or warranty of any kind, be it expressed or implied, as to the accuracy, completeness, or validity of the information. Any performance parameters, geometric tolerances, specific design features, quality and types of materials, or processes should not be inferred to represent what will be delivered by third-party suppliers or manufacturers through Xometry’s network. Buyers seeking quotes for parts are responsible for defining the specific requirements for those parts. Please refer to our terms and conditions for more information.

What is Laser Welding and How Does It Work?

Laser welding is not as prevalent as technologies like MIG, TIG and arc welding. That’s mainly because until recently, it required major investments.  

Around the year 2006 however, fiber laser technology made important advances, drastically bringing down costs. And for the last 18 years, further advances have continued to drive down production costs and the cost of ownership of laser welding machines. 

We’ve reached a time where laser welding is not only viable for high-volume production lines, but also for all types of manufacturers. We’re even seeing a rise in handheld devices . 

With laser welding being so accessible, now is a good time to explore this proven process that promises precision, speed, and cost effectiveness. 

Table of Contents

General Information

Conduction vs. keyhole welding, welding of dissimilar metals, adjustable ring modes, when was laser welding invented, what types of lasers can weld, what are the advantages of laser welding, what are the disadvantages of laser welding, battery welding machine, handheld system, robot welding machine (remote welding), robot welding of car frame, what is laser welding and how does it work.

Laser welding is a precise process that produces very little deformation compared to traditional welding methods. It uses a high-energy laser beam to fuse metals together, creating a strong metallurgical bond. As the energy from the laser beam is absorbed by the surface, the heat causes the surface to melt, forming a molten pool that resolidifies in a few milliseconds. 

The power density is very high, resulting in a concentrated heat source of millions of watts per cm 2 .  For a fast laser welding speed or a deep penetration, more laser power is needed. Laser power is the main factor that drives up the cost of a machine. 

Laser welding can be used on any material that can melt and resolidify. This means that it is not only used to weld metals like aluminum, copper, and stainless steel, but also other types of materials, including certain types of thermoplastics, glasses, and composites.

Image courtesy of Mammoth Memory .

Conduction welding (left) and Keyhole welding (right). Image courtesy of  The Fabricator .

The two main types of laser welding processes—conduction welding and keyhole welding—work differently.

Conduction welding is a soft process where the laser beam slowly melts the metal. During this type of welding, the metal’s temperature goes beyond its fusion point and achieves the liquid state but never goes into the gaseous state. Heat transfer within the metal is similar in all directions.  

Conduction welding is slower but generates higher-quality results with little or no spatter and low fumes. 

Keyhole welding is a fast but aggressive process that melts and vaporizes the metal, digging deeper into the material. The metal reaches its fusion temperature and even its vaporization temperature in some areas. As a result, part of the melting pool is in the gaseous state and can cause spatter. Heat transfer within the metal is mostly perpendicular to the laser beam.  

Keyhole welding is ideal for high-volume production lines because it is faster, but it can lead to porosity and a higher heat affected zone (HAZ). 

Laser welding of dissimilar metals is possible, but it is not always easy or feasible. Different metals have different fusion temperatures, absorb a different percentage of light, and conduct heat at different rates. 

When laser welding dissimilar metals, the two metals do not merge into a homogeneous mixture but rather join at the interface between the two metals. This creates a joint that is not as structurally strong as when welding the same metal.

To address this issue, two strategies are available:

• Filler material can be used to create stronger joints (like other welding techniques). In this case, we’re talking about a process called laser brazing—and not laser welding.
• The laser beam can be oscillated to help fuse the different metals more slowly. This process is called laser wobbling and requires additional optical components. Laser wobbling offers other benefits, as it helps get rid of gasses that would otherwise create porosity in the joints.

The first experiments with laser welding go back to the 1960s—shortly after Ted Maiman built the first laser. But it wasn’t until 1967, after researchers at the Battelle Memorial Institute did a demonstration of laser welding, that manufacturers began to see real potential for industrial applications. 

To understand how this process has come to hold such an important place in manufacturing, we need to go back to the invention of the laser itself. Here are key technological advancements that have shaped the development and adoption of laser welding as we know it today. 

• 1917 – Albert Einstein discovers stimulated emissions, providing the background knowledge needed to amplify light into laser beams.  
• 1957 – Gordon Gould develops the theoretical framework for the laser.  
• 1960 – Ted Maiman builds the first laser—a ruby laser— opening the door to potential applications. 
• 1960s – Various experiments are conducted to demonstrate the feasibility of laser welding.  
• 1962 – Researchers at the American Optical Company use an Nd:Glass laser to weld steel and titanium. 
• 1963 – Elias Snitzer demonstrates the first fiber laser, but it is limited in terms of output and efficiency compared to other lasers. 
• 1964 – At Bell Laboratories, Geusic et al. invent the Nd:YAG laser, which provides more power and efficiency than Nd:Glass lasers. 
• 1967 – Researchers demonstrate the practical applications and viability of laser welding at the Battelle Memorial Institute . This paves the way for further development and widespread adoption. 
• 1970 – At the Western Electric Company, CO2 lasers are used for laser welding for the first time, providing more power and lower costs than solid-state lasers like Nd:YAG lasers. 
• 1980s – Fiber lasers that provide higher beam quality and efficiency, lower maintenance, and easier integration are introduced at Southampton University in the UK. 
• 1990s – Laser systems begin to be integrated with robotic arms for automated welding processes. These first systems require that the workpiece be positioned close to the laser source. 
• 2000s – Advancements in fiber laser technology make laser welding affordable for a wider range of manufacturers. Advancements in scan heads pave the way for remote welding, making it possible to precisely direct laser beams from afar. 
• 2010s – Remote laser welding systems become increasingly feasible and commercially available, enabling the delivery of laser energy to the workpiece through fiber-optic cables over longer distances. 

Laser welding technology continues to evolve on a wide range of aspects, including in terms of laser power, optical components, beam quality, scanning heads, and computer control systems. 

Fiber lasers are the most prevalent types of lasers used for welding, but other types of lasers can be used as well, including blue lasers, green lasers, CO2 lasers, Nd:YAG lasers, and diode lasers.  

Let’s look at each type of laser to understand how they can be used for welding. 

When choosing a type of laser, one of the important factors to consider is its wavelength. Each type of metal absorbs and reflects wavelengths at varying percentages. If a wavelength is absorbed well, less laser power is needed.

The graph below provides an overview of common types of metals and their absorption spectrum for different wavelengths. 

Image courtesy of  Novika Solutions .

Laser welding offers a wide range of advantages compared to other methods like MIG, TIG, and arc welding.  Let’s look at the most important ones.

• The heat affected zone (HAZ) is smaller. The energy of the laser beam is focussed in a very small area and is moved as early as possible. With this level of precision and control, only the areas that need to be heated are heated. There is no unnecessary heat input.
• Parts maintain better mechanical properties. Due to the low heat input, there is less heat distortion and part warping. With other welding methods, excess heat degrades mechanical properties and often creates the need for straightening. This extra step is not required with laser. 
• Engineers can design lower weight parts. With other welding methods, mechanical engineers often address the excess heat input by designing parts with thicker materials. But with laser welding, since there is as little heating as possible, it is possible to go for thin materials, which helps minimize product weight and material costs. This is very significant for manufacturers involved in the aerospace and automotive industries, where reducing vehicle weight is a key objective to improve range. 
• Small components can be welded due to the high level of precision. This is especially relevant for electronic components, tab connections, and similar applications. 
• Laser welding is faster than other processes. Thanks to fiber laser technology, industrial lasers can easily operate at several thousands of watts, which is more than enough to meet the most demanding production requirements. 
• Higher-quality welds are direct results of a better control over the process. During welding, the rapid heating and cooling of the material helps prevent quality issues. For example, there is a reduced likelihood of hydrogen embrittlement. Hydrogen embrittlement occurs when hydrogen penetrates the metal, causing mechanical damage. Due to the speed of the process, there is minimal time for hydrogen absorption and diffusion. 
• Laser welding is easy to automate due to factors like remote capabilities, minimal wear and tear, and repeatability. This makes it an interesting technology for manufacturers who have trouble finding specialized welders. 

There are not many disadvantages to laser welding, but they are still important to consider and address. Let’s look at them more closely. 

Laser safety is a serious issue during welding. The laser beam and its reflection can cause eye injuries, skin burns, and fire hazards. Ideally, the laser is enclosed in a class-1 laser safety enclosure that contains the laser beam and its reflections.  

For some applications, this can cause headaches. Large parts and structures such as ships can be difficult to contain in an enclosure. Other solutions than enclosures can be explored to contain the beam (for example, some solutions use clamping tools to block the beam).  

Handheld devices can be dangerous for operators who need to wear PPE and follow laser safety control measures. 

The initial investment can be pretty high too. Even if their cost keeps going down, lasers are still more expensive to acquire than alternatives. Add to this that most manufacturers are seeking automated solutions, and you’re looking at a serious investment.  

Examples of Laser Welding Machines

This battery laser machine is a fully enclosed solution that overcomes the challenges of welding batteries such as clamping adaptability, quality, and speed. It uses SCARA robots to perform clamping with high speed and precision.

This handheld system is an example of how laser welding products are becoming more accessible to a wider range of manufacturers. 

This machine demonstrates the ease of automation of laser welding, featuring remote welding, rotating fixtures, and robot handling of the laser head.

For large parts like car frames, enclosing the welding process can be problematic in terms of space usage and cost. This machine addresses this problem by enclosing the beam and its reflections with a well-designed clamping tool. This provides complete laser safety without the need for an enclosure.

With its precision, speed, and high level of control, laser welding offers amazing possibilities. For manufacturers new to this technology, it’s important to discuss your application with an expert who can help you:

• Rethink your product with laser welding in mind 
• Evaluate if laser welding is a good solution for you 
• Perform feasibility studies 
• Understand integration implications and costs

If you have a laser welding application for batteries, contact a Laserax expert to discuss your needs. 

Let Us Know Your Application

Contributor

Patrick Martel

Catherine Veilleux

Thank you for visiting nature.com. You are using a browser version with limited support for CSS. To obtain the best experience, we recommend you use a more up to date browser (or turn off compatibility mode in Internet Explorer). In the meantime, to ensure continued support, we are displaying the site without styles and JavaScript.

• View all journals
• Explore content
• About the journal
• Publish with us
• Sign up for alerts
• Open access
• Published: 03 October 2022

Inhibiting weld cracking in high-strength aluminium alloys

• Yanan Hu 1 , 2 ,
• Shengchuan Wu   ORCID: orcid.org/0000-0002-7437-2021 1 , 3 ,
• Zhao Shen   ORCID: orcid.org/0000-0001-7432-4604 5 , 6 ,
• Alexander M. Korsunsky   ORCID: orcid.org/0000-0002-3558-5198 7 ,
• Yukuang Yu 1 ,
• Xu Zhang   ORCID: orcid.org/0000-0001-8481-0059 2 ,
• Yanan Fu 8 ,
• Zhigang Che 9 ,
• Tiqiao Xiao 8 ,
• Sergio Lozano-Perez   ORCID: orcid.org/0000-0003-3387-5973 6 ,
• Qingxi Yuan   ORCID: orcid.org/0000-0002-2997-7654 10 ,
• Xiangli Zhong   ORCID: orcid.org/0000-0003-4772-1158 3 ,
• Xiaoqin Zeng 5 ,
• Guozheng Kang 1 , 2 &
• Philip J. Withers   ORCID: orcid.org/0000-0002-6896-0839 3  

Nature Communications volume  13 , Article number:  5816 ( 2022 ) Cite this article

11k Accesses

55 Citations

7 Altmetric

Metrics details

• Aerospace engineering
• Metals and alloys

Cracking from a fine equiaxed zone (FQZ), often just tens of microns across, plagues the welding of 7000 series aluminum alloys. Using a multiscale correlative methodology, from the millimeter scale to the nanoscale, we shed light on the strengthening mechanisms and the resulting intergranular failure at the FQZ. We show that intergranular AlCuMg phases give rise to cracking by micro-void nucleation and subsequent link-up due to the plastic incompatibility between the hard phases and soft (low precipitate density) grain interiors in the FQZ. To mitigate this, we propose a hybrid welding strategy exploiting laser beam oscillation and a pulsed magnetic field. This achieves a wavy and interrupted FQZ along with a higher precipitate density, thereby considerably increasing tensile strength over conventionally hybrid welded butt joints, and even friction stir welds.

Similar content being viewed by others

High-strength and crack-free welding of 2024 aluminium alloy via Zr-core-Al-shell wire

In situ observation and reduction of hot-cracks in laser additive manufacturing

Flat hardness distribution in AA6061 joints by linear friction welding

Introduction.

Modern welding technology can trace its roots as far back as the latter half of the 19th century 1 . Nowadays, it is an everyday tool in the energy, shipbuilding, automotive, aircraft, aerospace, and railway industries. It enables the assembly of lightweight structures, which is of paramount importance in reducing energy consumption and carbon emissions 2 . In this respect, lightweight aluminum (Al) alloys have been increasingly deployed in recent decades. Use of high-strength Al alloys, such as aluminum–lithium (Al–Li) and 7000-series (Al–Zn–Mg–Cu) alloys, in particular, has become increasingly widespread 2 , 3 . One longstanding challenge is to overcome the local softening and cracking issues associated with conventional welding. This has seriously hampered long-term service applications 4 , 5 , 6 and led to a focus on solid state friction stir welding 7 . In many cases, this strength reduction is associated with the so-called fine equiaxed zone (FQZ) prevalent in fusion welds of these materials. At the microscopic level, the unique microstructural features associated with the FQZ are key because the precipitation characteristics have a considerable influence on the mechanical properties and failure behavior. FQZs have also been observed in the welded joints of other Al alloys and steels (see Supplementary Table  1 ). To date, issues associated with FQZs have not received the attention that they deserve from a structural integrity point of view and a better understanding of the softening and failure mechanisms related to the FQZ is required.

The formation of the FQZ has been well studied in terms of the alloy constituents, solidification, base materials (BM), welding parameters, thermal history, and molten pool dynamics 6 , 8 . For Al–Li and Zr-containing Al alloys, the FQZ is formed through heterogeneous grain nucleation aided by Al 3 (Li x , Zr 1-x ) and Al 3 Zr respectively 6 . However, the details of the softening and cracking behavior of the FQZ have not yet been clearly elucidated. This may be attributed to the fact that the FQZ is very narrow, containing very fine equiaxed grains posing a challenge to the precise characterization of the microstructure, properties, and the associated damage accumulation sequence. Further, the in-service performance largely depends on microstructural features across length scales ranging from the macro- to the nano-scale. Recently, there has been a focus on combining various imaging methods with different resolutions to link the micro- and nano-scale features through what is termed ‘correlative characterization’ approaches 9 . Here the behavior of the FQZ is interrogated across multiple length scales by multiscale correlative tomography 10 , 11 , 12 to shed light on the various damage evolution mechanisms and their sequence.

Here we consider the FQZ arising from a hybrid laser and arc butt weld (HLAW) of 7050 Al alloy (see Fig.  1a ) bringing together multiple techniques to study the microstructure across the scales (Fig.  1b–f ). The relatively high rates of crystal nucleation and the very high solidification rates near the fusion boundary (FB) give rise to a large number of fine non-dendritic equiaxed grains located between the heat affected zone (HAZ) and the central weld metal (WM)—see Fig.  1d . The narrow (50–100 μm wide) FQZ is frequently missed when mapping hardness across the weld zone by Vickers hardness testing. Here the grain orientation, size distribution, and grain boundary characteristics in the FQZ have been characterized by electron backscatter diffraction (EBSD). These grains exhibit a random crystal orientation and their equivalent diameters range from 3 to 10 μm (average size ~7 μm). Our EBSD measurements show that in the FQZ a large fraction (~87%) of the boundaries are high-angle grain boundaries (HAGBs, >10°). Electron probe micro analysis (EPMA) reveals that these grain boundaries are highly enriched by the strengthening elements Zn, Mg, and Cu due to segregation during solidification (Fig.  1e ). As a result, the grain boundaries are decorated with interconnected phases. Furthermore, the combination of severe elemental segregation and fast cooling act to limit the reprecipitation of the precipitates within the grains. Transmission electron microscopy (TEM) clearly shows that the distribution of the precipitates in the FQZ (Fig.  1f ) represents a volume fraction of just ~0.2% and an average radius of ~11 nm, compared to ~3.7% and ~15 nm respectively in the BM.

a Schematic of the welding process where the HAZ, WM and FB represent heat affected zone, weld metal and fusion boundary, respectively. b Schematic of in situ tensile synchrotron radiation X-ray micro computed tomography (microCT) at the 13HB beam line (BL13HB) of the Shanghai Synchrotron Radiation Facility (SSRF). c High-resolution synchrotron X-ray nanoCT at the 4W1A beam line (BL14W1A) of the Beijing Synchrotron Radiation Facility (BSRF). d Electron backscatter diffraction (EBSD) inverse-pole figure (IPF) map across the fusion boundary (HAZ (left); FQZ (centre), weld (right)) where the high-angle grain boundaries (HAGBs, >10°) and the low-angle grain boundaries (LAGBs, 5–10°) are colored black and red, respectively. e Electron probe micro analysis (EPMA) maps showing chemical constitutions of the intergranular phases. f Bright-field transmission electron microscopy (TEM) images of precipitates in the interior of FQZ grains at different magnifications.

In this work we focus our attention on identifying the critical microstructural aspects affecting the local softening and intergranular failure to better understand the effect of the FQZ on softening and failure. First, we estimate the contribution to the yield strength of the grain size, dislocation density, solute and precipitate strengthening mechanisms using classical strengthening models across the different regions of the weld. Second, we employ a multiscale correlative tomography procedure to investigate the damage evolution and intergranular failure behavior, through in situ synchrotron radiation X-ray micro computed tomography (SR-μCT) during tensile straining (Fig.  1b ), high-resolution synchrotron X-ray nanoCT (Fig.  1c ) and energy dispersive spectrometry (EDS) in the TEM. Finally, we have developed an effective strategy to mitigate the effects of the FQZ by oscillating the laser beam and applying a pulsed magnetic field to disturb the FQZ, leading to a higher tensile strength, even reaching the strength level obtained by solid state welding.

Softening mechanism

Nanoindentation testing (Fig.  2a ) shows the narrow FQZ to be, by some margin, the softest region across the weld zone (hardness ~54% that of the BM). To understand the origin of the softening, we first performed a quantitative analysis of the average precipitate characteristics (radius, r , and volume fraction, f ), grain sizes ( d ), solute concentrations ( c Zn , c Mg , and c Cu ), and dislocation densities ( ρ ) across the four regions of the weld (see below). The results are summarized in Table  1 . Based on these values the various strengthening contributions (grain size strengthening (Δ σ gb ), solid-solution strengthening (Δ σ ss ), dislocation strengthening, (Δ σ dis ), and precipitation strengthening (Δ σ ppt )) can be estimated by the Hall-Petch, Fleischer, Bailey-Hirsch, and Orowan models, respectively (see Methods).

a The average nanoindentation hardness recorded for each region across the weld where the BM, HAZ, WM and FQZ represent base metal, heat affected zone, weld metal and fine equiaxed zone, respectively. b The estimated strengthening contributions arising from the precipitate (Δ σ ppt ), grain size (Δ σ gb ), solute (Δ σ ss ) and dislocation (Δ σ dis ) strengthening, respectively. σ 0 represents the baseline strength of pure aluminum. c (From left to right) distributions of precipitates in the grain interiors, distributions of grain sizes, scanning points used for chemical analysis by electron probe micro analysis (EPMA) and density of geometrically necessary dislocations ( ρ GND ).

Conventionally, arithmetic addition (Eq. 1 ) and quadratic addition (Eq.  2 ) have most commonly been used to combine these strengthening contributions 13 , 14 :

where σ 0 represents the baseline strength of pure aluminum ( σ 0  = ~10 MPa 15 ). We have considered both models to predict the yield strength of the BM giving strengths of ∼ 450 MPa and ∼ 430 MPa respectively.

Given the former is simpler, closer to the experimental value ( ∼ 451 MPa in Supplementary Fig. 1 ) and because the main focus is on evaluating the relative importance of the individual mechanisms rather than finding a model that precisely matches the yield strength, we have used arithmetic addition to understand the relative contributions in each region of the weld. It has been used in many investigations 16 , 17 , 18 . The relative contributions are shown in Fig.  2b and in Supplementary Table  2 .

Irrespective of the addition rule, it is evident that precipitation strengthening is the most significant contributor to the strength in the BM while grain size strengthening is relatively insignificant. Although the grain refinement in the FQZ means that the grain boundary strengthening is ∼ 2 times larger, the loss in strength arising from the reduction in the precipitation strengthening (a quarter that for the BM) means that overall it is significantly softer.

It has been suggested that the evaporative loss of Zn and the inverse segregation of Cu are the main reasons for the low level of strengthening precipitates in the WM 19 . For the FQZ, the peak temperature is significantly lower than the peak temperature in the WM, so that significant evaporative loss of Zn is unlikely. Furthermore, the laminar boundary layer in the FQZ may suppress the inverse segregation of Cu 20 . As a result, the number of precipitates within the grains in the FQZ, which is much smaller than that for the WM (Fig.  2c ), most probably arises from a combination of the extensive segregation of the strengthening elements to the very many grain boundaries (Fig.  1e ) and the fast solidification together which limit the extent of reprecipitation of the precipitates on cooling.

Damage evolution mechanism

The sharp variation in mechanical properties at the FQZ would be expected to lead to inhomogeneous plastic deformation and a high cracking sensitivity. An image-based 3D finite element (FE) simulation undertaken in ABAQUS confirms that, even at low stresses, plastic strain concentrates in the lower parts of the FQZ and the WM due to their low strengths (Fig.  3a ). This explains why a crack is observed by time-lapse microCT (Fig.  3b ) to initiate from the weld toe and then propagate approximately along the FB with increased loading ( Supplementary Movie and Supplementary Fig.  2 ). Post-mortem fractography (Fig.  3c ) reveals the fracture surface to be relatively flat exhibiting smooth curving facets. The size of these curved facets is consistent with the equiaxed grains inside the FQZ. This suggests that the crack grows primarily along the grain booundaries in the FQZ (Fig.  3c and Supplementary Fig.  3 ).

a Distribution of equivalent plastic strain (PEEQ) across the weld zone predicted by an image-based 3D finite element (FE) simulation where the applied stress σ  = 60 MPa perpendicular to the weld. b 3D volume renderings (pores rendered green and cracks yellow) acquired at two loading stages by in situ tensile synchrotron micro computed tomography (microCT) to determine the preferential damage nucleation site and damage evolution. c Post-mortem fractography of a failed welded joint by scanning electron microscope (SEM). d NanoCT volume renderings of the spatial distribution of intergranular phases prior to deformation. e Quantitative analysis of the projected areas of nucleated micro-voids when σ  = 270 MPa. f Spatial distribution of intergranular phases and nucleated micro-voids by high-resolution synchrotron X-ray nanoCT when σ  = 270 MPa showing the large-sized long-range connected voids (green), the intergranular phases (yellow) and the nucleated micro-voids (red). g NanoCT visualization of the interaction between the intergranular phases (yellow) and nucleated micro-voids (red). h Intergranular phase inducing micro-void nucleation observed by (transmission electron microscopy-energy dispersive spectrometry) TEM-EDS, showing TEM image, selected area electron diffraction (SAED) and EDS spectra.

In order to visualize the 3D nature of the intergranular phases, a micropillar was excised from a region of FQZ by plasma FIB for an unstrained sample for examination by synchrotron X-ray nanoCT (Fig.  1c ). The intergranular phases (rendered yellow in Supplementary Fig.  4 and Fig.  3d ) include rod-shaped, network-shaped and dendritic morphologies. The different projections show the morphology of the intergranular phases to be anisotropic (Fig.  3e ); presenting a network-like morphology on both the x-y and z-x planes, and a flake shape on the y-z plane, in which the y-z plane is parallel to the plane consisting of the weld thickness and welding direction. This indicates that the intergranular phases tend to occupy the spaces between the grains along the weld thickness.

In order to investigate the damage evolution micro-mechanism in greater detail, two micropillars (40 μm in diameter) were excised by plasma focused ion beam (PFIB) from the weld when σ  = 270 MPa, as indicated in Fig.  3b , to perform synchrotron X-ray nanoCT observations (Fig.  3f and Supplementary Fig.  4 ). These show a large interconnected void (colored green) in the upper part and many isolated micro-voids (colored red) in the bottom half of the two micropillars, both types having initiated at the intergranular phase. The micro-voids are randomly distributed irregular, ellipsoidal or spherical shapes. Quantitative analysis shows that the equivalent diameter of these micro-voids is 100-500 nm, with an average value of ~240 nm, with those below 300 nm representing 90% of the distribution. Their projected area on the z-x plane is generally smaller than that on the x-y and y-z planes (Fig.  3e ) suggesting they nucleate and grow along the welding direction perpendicular to the loading direction.

It is evident from Fig.  3g that all the micro-voids (red) nucleate from the intergranular phases (yellow) due to the plastic incompatibility. They form either within the intergranular phases (indicated by A) by phase fracture, or adjacent to it (indicated by B) suggesting debonding of the particle-grain boundary interface, or, most commonly, at the grain boundary triple junctions (indicated by C and D).

To identify the composition of the intergranular phases a typical interphase region was excised by PFIB for TEM analysis. A number of intergranular phase regions were located in bright field imaging and identified through a combination of SAED and EDS analysis. Representative results are shown in Fig.  3h and Supplementary Fig.  5 which are consistent with AlCuMg. This is not unexpected, brittle AlCuMg phases are often observed in Al alloys with higher Cu content, such as the 2000-series Al-Cu alloy. Here we have very high Cu concentrations due to the segregation and this explains the presence of AlCuMg phases. The plastic mismatch between these hard regions and the soft grain interiors promotes void nucleation at the grain boundaries (Fig.  3 and Supplementary Fig.  6 ).

Unsurprisingly, the high resolution EBSD based geometrically necessary dislocation (GND) density maps in Supplementary Fig.  7 show that the GND density in the FQZ has increased significantly when the applied stress reaches 320 MPa (slightly above the yield strength of the FQZ in Supplementary Table  2 ). It is noteworthy that the increased GND density is not concentrated at the grain boundaries which serve as strong barriers for dislocation movement, but rather they are relatively homogeneously distributed in the grain interiors and at the grain boundaries. This is also confirmed by EBSD measurements recorded in situ during tensile straining (see Supplementary Fig.  8 ).

A hybrid welding technique

Clearly the FQZ poses a significant threat to the reliable in-service performance of welded structures, however, research indicates that it cannot be entirely eliminated by varying the welding parameters 6 . A key question therefore is, how can we re-engineer the FQZ to mitigate its deleterious effect on weld performance? The first strategy to consider is post-weld heat treatment (PWHT). An in situ heating EBSD experiment shows that the PWHT (470 °C/30 min) had essentially no effect on the FQZ, for the surface grains at least (Supplementary Fig.  9 ). This observation is supported by destructive observations of the FQZ and the fact that after PWHT the yield behavior (ultimate tensile and yield strengths of σ b  = 375 ± 6 MPa and σ p0.2  = 295 ± 5 MPa) was similar to that ( σ b  = 406 ± 11 MPa and σ p0.2  = 292 ± 10 MPa) of the as-welded joints.

An alternative strategy is to enhance molten pool turbulence so as to modify the weld microstructure 21 , 22 . Here we investigated the effect of oscillating the laser beam during HLAW and applying a pulsed magnetic field (OLHW + m) immediately following welding after the solid has formed. In this way we have produced butt-welded joints having the wavy morphology shown in Fig.  4a . At the microscale, the EBSD IPF maps in Fig.  4d show that the FQZ at the fusion boundary has been broken up. This is evident in Fig.  4d where it is alternately arranged with coarse equiaxed dendritic structures on the z-x section and intermittently distributed near the FB on the y-z section. Its curved nature can also be seen in the x-y section.

a Morphology of the lower weld surface using ultra-depth 3D microscopy. b Engineering stress-strain curves of traditional gas metal arc weld (GMAW) joints, HLAW joints and OSHW + m joints. c Fatigue crack growth rate curves (d a /d N -Δ K where d a /d N is the fatigue crack growth rate, a is the crack length, N is the number of loading cycles, Δ K is the stress intensity factor range) of HLAW joints and OSHW + m joints with side views of the macro failure path. The applied stress σ is perpendicular to the weld. d (From left to right) Electron backscatter diffraction (EBSD) X inverse-pole figure (IPF) maps on y-z, z-x (where the weld root is located) and x-y sections of OSHW + m joints. The WM, FQZ and HAZ represent weld metal, fine equiaxed zone and heat affected zone.

The tensile strength of the OSHW + m weld (~470 MPa) is clearly significantly (20%) better than the conventional HLAW (Fig.  4b ) and unsurprisingly, far superior (60% higher) than that made by gas metal arc welding (GMAW). For comparison it is slightly greater than for welds made by friction stir welding (FSW) ( σ b  = 450 MPa 23 ) after the same natural aging treatment (~3000 h). The fracture section in Fig.  4d (right) shows that the crack propagates initially the soft FQZ but then grows into the weld zone despite its higher strength. This is because in this region the primary crack has to deflect to continue to propagate due to the curved and discontinuous nature of the FQZ. As a result, the gradient structure (mixture of FQZ and WM) near the fusion boundary for the OSHW + m weld displays a higher cracking resistance compared to conventional HLAW having a straight FQZ 24 .

Given that the loss of precipitate strengthening in the FQZ is critical to the extensive softening of the FQZ it is important to compare the precipitate microstructures for the modified and conventional hybrid laser welds (Supplementary Fig.  10 ). It is evident that the number of precipitates in the FQZ is much higher than for the conventional weld, with volume fractions of ~0.9% and ~0.2%, respectively. This is because the laser beam oscillation more effectively controls the distribution of alloying elements in the weld and mitigates macro-segregation 25 . This retains the elements in solution leading to their subsequent reprecipitation as strengthening precipitates in the interior of the grains and consequently higher strength.

As a result of the FQZ morphology, the crack follows a more tortuous macroscopic failure path in comparison with the HLAW joints leading to a relatively slower crack growth rate Fig.  4c , especially in the microstructure sensitive near-threshold and unstable crack growth regions. The crack deflection gives rise to several periods of arrested growth, probably due to the crack growing through a mixture of a harder phase (HAZ and WM) and a softer phase (FQZ) regions near the fusion boundary 26 , 27 .

In summary, we have quantified the strengthening mechanisms giving rise to significant softening in the FQZ and delineated the sequence by which the crack initiates intergranularly due to the presence of AlCuMg phases. To counter the deleterious effect of the FQZ on the low strength of hybrid laser welded 7050 aluminum alloys, we have introduced an oscillated laser hybrid weld with externally applied magnetic field. This process disrupts the FQZ distribution in three dimensions and increases the precipitate density. Together these changes radically increase the tensile strength extending the UTS to ~470 MPa. This is ~90% of that of the base metal (UTS = 521 MPa), comparing favorably with the strength of the solid state friction stir welded joints 28 , 29 .

Materials and welding procedure

A hybrid fiber laser-pulsed arc welding system was employed to produce 2-mm-thick 7050 Al alloy butt-welded joints, where the weld was perpendicular to the rolling direction. The filler material was ER5356 Al-Mg wire having a diameter of 1.2 mm. The welding parameters were: laser power P  = 3 kW, electric current I  = 100 A, welding speed v  = 6 m/min and defocusing distance Δ = −1 mm.

A second HLAW experiment (designated OSHW + m) was undertaken but employing laser beam oscillation following a saw-tooth trajectory down the weld. The optimized welding parameters were: laser power, 6 kW; welding speed, 8 m/min; defocusing distance, +2 mm; oscillating diameter, 1-3 mm; and oscillating frequency, 100-300 Hz. Post-weld electropulsing treatment was then applied to the OSHW joints immediately following welding. Electropulsing was achieved by passing and electrical current through a material using a home-made device equipped with a capacitor bank discharge circuit (Fig.  1a ). We chose a discharge voltage of 4 V, an alternating current pulse with duration 100 ms, followed by a 10 s natural air cooling.

Microstructural characterization

For EBSD analysis, the samples were polished with colloidal silica and milled by Ar ion beam. The crystallographic data was estimated using a TESCAN MIRA3 SEM equipped with a Bruker e-Flash FS detector. In order to calculate the GND density, cross-correlation based EBSD method was adopted. In the current work, 12-bit 1200 × 1200 pixel patterns were saved to hard disk during EBSD scan. The calculated rotation gradients could be linked to Nye’s dislocation tensor 30 and the dislocation density was estimated using L1 optimization by minimizing the total dislocation line energy. A full description of the method used for GND calculation can be found in  Supplementary material 31 .

The samples for EBSD analysis were further analyzed to determine the variation in alloying constituents near the fusion boundary and the elemental distributions across the grains and the grain boundaries using a JEOL JXA-8230 EPMA. The content of the solute elements in the interior of the grains was also quantitatively analyzed. Both the size and volume fraction of the precipitates within the grains were examined by a Tecnai G2 F30 S-TWIN TEM on the 20-μm-thick samples prepared by an Ion Beam Thinner. The chemical maps were acquired by EDS using four Bruker SDD detectors.

Mechanical testing

A G200 KEYSIGHT Nano Indenter tester was used to perform nanoindentation hardness testing. The sample surface was carefully polished to achieve a surface roughness a quarter of the maximum indentation depth 32 . The maximum indentation depth was 1000 nm. The Oliver-Pharr method was adopted to calculate the nanohardness of different zones of the joints 33 .

Monotonic tensile testing was conducted on three kinds of welded joints manufactured by GMAW, HLAW and OSHW + m, with a minimum width of 12 mm, a gage length of 65 mm and a thickness of 2 mm. The nominal strain rate was 1.0 mm/min. The loading direction was perpendicular to the welding direction. An extensometer with a gauge length of 20 mm was used to determine the yield stress and elastic modulus.

Fatigue crack growth rate testing was conducted on compact tension (CT) specimens cut from both HLAW and OSHW + m joints, using a frequency of 10 Hz and a load ratio of 0.1. The thickness and width of the CT specimens were 2 mm and 48 mm, respectively. A 10‐mm long pre‐crack was emanated from the notch tip along the weld. The crack growth rate (d a /d N ) was determined by the 7‐point increment polynomial method.

Multiscale correlative tomography

We have threaded together in situ X-ray micro tomography, X-ray nano tomography and TEM-EDS characterization all for the same region. In situ tensile SR-μCT was performed on a specimen with an area of minimum section of 2 mm 2 . The role of the large volume in situ tensile SR-μCT was to provide a means for identifying and locating RoI containing interesting features for higher resolution X-ray nano-tomography. A FEI Helios PFIB system 34 based on Xe + was used to extract a pillar of approximately 40 μm diameter and 40 μm height, at the location of the RoI, for nanoCT. To characterize the spatial distribution of the nucleated micro-voids and intergranular phases, a nanoscale X-ray CT experiment was performed on the BL4W1A at the Beijing Synchrotron Radiation Facility (BSRF), China. The ‘large field of view’ mode was adopted with a field-of-view size of 60 μm × 60 μm, a pixel size of 64.1 nm, a resolution of 100 nm, an energy of 8 keV and an exposure time of 12 s. All visualization was performed using the Avizo software package. To further determine the critical intergranular phases inducing the micro-void nucleation, thin slices of the RoI were removed for analysis in the TEM using a Tecnai G2 F30 S-TWIN TEM.

X-ray CT Image-based finite element modeling

The volumetric image datasets of metallurgical defects from SR-μCT were segmented using the tomography software Avizo. Considering the meshing process and computing feasibility, the sharp variations of defects were smoothed by adjusting the smooth factor. Avizo-based 3D reconstruction was able to directly produce surface meshes (linear triangles) of inner defects and outer surfaces of the samples. Then the volume filling of the samples was conducted using the software HyperMesh, producing over 2,657,544 linear tetrahedral elements (C3D4) for FE analysis. Both linear elastic and elasto-plastic responses were considered, and the Young’s modulus and Poisson’s ratio were 70 GPa and 0.33, respectively. The bilinear isotropic model was employed for the elastic-plastic prediction where the stress–strain curves were determined by the load-displacement curves recorded during nanoindentation testing using the reverse analysis based on dimensionless functions. The right surface in Fig.  3a was fixed in all directions and a remote stress of 60 MPa in y direction was applied on the left surface.

Crystal plasticity finite element modelling

The modeling was performed on a model with a dimension of 6.4 μm × 4.29 μm × 0.19 μm, consisting of eight grains and 64,437 C3D8R elements using ABAQUS/CAE. The grain orientations were assigned in terms of three Euler angles, {θ, φ, Ω}, representing rotations from the crystal basis to the global basis, according to the EBSD results. The morphology of intergranular phases was referred to the bright-field TEM images. The boundary conditions are: on the x  = 0 surface, the displacement in the x direction is zero ( u x  = 0); on the y  = 0 surface, u y  = 0; on the z  = 0 surface, u z  = 0; on the y  = 6.4 μm and x  = 0.19 μm surfaces, the traction is zero; on the z  = 4.29 μm surface, the velocity in the z direction is constant, corresponding to a strain rate of 4 × 10 −5  s −1 . Each intergranular phase is regarded as a rigid body through the rigid body constraint. The face-centered cubic structure of α-Al matrix is anisotropic and defined by three elastic constants C 11 , C 12 , and C 44 , where C 11  = 106.43 GPa, C 12  = 60.35 GPa, and C 44  = 28.21 GPa 35 .

The nonlocal crystal plastic constitutive laws in the DAMASK simulation platform developed by the Max-Planck-Institut für Eisenforschung in Germany were adopted 36 . The modelling is based on the dislocation mechanism. The dislocation flow term is derived from the relationship between the plastic strain gradient and the geometrically necessary dislocation density. The dislocation density evolution includes dislocation generation, dislocation annihilation, dislocation pair formation and annihilation, and dislocation flow between material points. The plastic slip rate is described by the Orowan equation:

where ρ α mobile is mobile dislocation density, v α is the dislocation velocity and the value of the Burgers vector b is 2.48 × 10 −10  m.

Furthermore, the dislocation multiplication can be expressed as:

where \({\dot{\gamma }}_{{e}^{+}}^{\alpha }\) , \({\dot{\gamma }}_{{e}^{-}}^{\alpha }\) , \({\dot{\gamma }}_{{s}^{+}}^{\alpha }\) and \({\dot{\gamma }}_{{s}^{-}}^{\alpha }\) are the plastic shear rate for different types of dislocation. k 1 and k 2 are the proportion coefficient. λ α is the mean free path of dislocation. The contribution coefficient, k 1 , and multiplication coefficient, k 2 , are k 1  = 0.1 and k 2  = 10.

However, the movement of dislocation on the slip plane is hindered by the forest dislocation. The effective shear stress on the slip plane is:

where τ α is the resolved shear stress along the slip plane of the second Piola-Kirchhoff stress.

In addition, the interaction strength between dislocations is:

where ξ αα ׳ is the dislocation interaction coefficient among slip systems. G is the shear modulus.

The relationship between the plastic distortion tensor β α and the dislocation density is associated by the Nye tensor α α :

where l α represents the unit vector along the dislocation line on the α-slip system.

By defining the dislocation flux f α as the product of the dislocation density and the dislocation velocity, the flow relationship of dislocations between material points can be derived:

Estimating the strengthening contributions

The yield strength increase arising from the Orowan by-passing mechanism can be expressed as follows 37 :

where M  = 3.06 is the Taylor factor of face-centered-cubic (fcc) alloy. G  = 27 GPa is the shear modulus of aluminum alloy. b  = 0.286 nm is the value of the Burgers vector of Al alloy. v  = 0.33 is the Poisson’s ratio of aluminum alloy. \(\overline{r}=\sqrt{2/3r}\) in which r is the mean precipitates radius. \({L}_{{{{{{\rm{p}}}}}}}=2\overline{r}(\sqrt{\pi /4f}-1)\) is the inter-precipitate distance in which f is the precipitates volume fraction. In our case, both r and f are extracted from the TEM images using image processing software ImageJ.

The strengthening contribution arising from the grain boundaries can be calculated using the classical Hall-Petch equation 38 :

where k is the strengthening coefficient specific to each material and k  = 0.12 for aluminum alloy. d is the grain size derived from on the EBSD analysis.

The strengthening contribution arising from solute strengthening has been estimated using the Fleischer equation 39 :

where Δ σ i is the theoretical strengthening efficiency for each element (Δ σ Zn  = 2.9 MPa/wt. %, Δ σ Mg  = 18.6 MPa/wt. % and Δ σ Cu  = 13.8 MPa/wt. % 39 ). c i is the concentration of the solute element (in wt. %), which is measured by EPMA in the study.

The Bailey-Hirsch equation 40 is adopted to evaluate the dislocation strengthening due to the interaction of dislocations:

where α is a material constant ( α  = 0.2 of the fcc alloy). ρ GND is the GND density, which is determined by EBSD analysis.

Data availability

The experimental data that support the findings of this study are available from the corresponding authors upon request. The source data underlying Figs.  1c , 2c , 3a, b, d–f , 4b , c are available in the database under accession code 0101 [ https://pan.baidu.com/s/12nfseoV38kelFSYTZ_8k5Q ].

David, S. A. & DebRoy, T. Current issues and problems in welding science. Science 257 , 497–502 (1992).

Article   ADS   CAS   Google Scholar  

Sokoluk, M., Cao, C. Z., Pan, S. H. & Li, X. C. Nanoparticle-enabled phase control for arc welding of unweldable aluminum alloy 7075. Nat. Commun. 10 , 98 (2019).

Article   ADS   Google Scholar  

Dursun, T. & Soutis, C. Recent developments in advanced aircraft aluminum alloys. Mater. Des. 56 , 862–871 (2014).

Article   CAS   Google Scholar  

Zhang, X. Y., Yang, W. X. & Xiao, R. S. Microstructure and mechanical properties of laser beam welded Al-Li alloy 2060 with Al–Mg filler wire. Mater. Des. 88 , 446–450 (2015).

Fu, B. L. et al. Microstructure and mechanical properties of newly developed aluminum-lithium alloy 2A97 welded by fiber laser. Mater. Sci. Eng. A 617 , 1–11 (2014).

Gutierrez, A. & Lippold, J. C. A proposed mechanism for equiaxed grain formation along the fusion boundary in aluminum-copper-lithium alloys. Weld. J. 77 , 123–132 (1998).

Google Scholar  

Threadgill, P. L., Leonard, A. J., Shercliff, H. R. & Withers, P. J. Friction stir welding of aluminium alloys. Int. Mater. Rev. 54 , 50–93 (2009).

Article   Google Scholar  

Dev, S., Murty, B. S. & Rao, K. P. Effects of base and filler chemistry and weld techniques on equiaxed zone formation in Al-Zn-Mg alloy welds. Sci. Technol. Weld. Join. 12 , 598–606 (2008).

Burnett, T. L. & Withers, P. J. Completing the picture through correlative characterization. Nat. Mater. 18 , 1041–1049 (2019).

Daly, M. et al. A multi-scale correlative investigation of ductile fracture. Acta Mater. 130 , 56–68 (2017).

Niverty, S., Kale, C., Solanki, K. N. & Chawla, N. Multiscale investigation of corrosion damage initiation and propagation in AA7075-T651 alloy using correlative microscopy. Corros. Sci. 185 , 109429 (2021).

Burnett, T. L. et al. Correlative Tomography. Sci. Rep. 4 , 4177 (2014).

Fang, H. J. et al. Evolution of texture, microstructure, tensile strength and corrosion properties of annealed Al-Mg-Sc-Zr alloys. Mater. Sci. Eng. A 804 , 140682 (2021).

Ferguson, J. B. et al. On the superposition of strengthening mechanisms in dispersion strengthened alloys and metal-matrix nanocomposites: Considerations of stress and energy. Met. Mater. Int. 20 , 375–388 (2014).

Lee, S. H. et al. Precipitation strengthening in naturally aged Al-Zn-Mg-Cu alloy. Mater. Sci. Eng. A 803 , 140719 (2021).

Wang, J. et al. Microstructure evolution and mechanical properties of the electron-beam welded joints of cast Al-Cu-Mg-Ag alloy. Mater. Sci. Eng. A 801 , 140363 (2021).

Masuda, T. et al. Achiening highly strengthened Al-Cu-Mg alloy by grain refinenment and grain boundary segregation. Mater. Sci. Eng. A 793 , 139668 (2020).

Sunde, J. K. et al. Linking mechanical properties to precipitate microstructure in three Al-Mg-Si(-Cu) alloys. Mater. Sci. Eng. A 807 , 140862 (2021).

Wu, S. C. et al. Porosity, element loss, and strength model on softening behavior of hybrid laser arc welded Al-Zn-Mg-Cu alloy with synchrotron radiation analysis. Weld. J. 92 , 64S–71S (2013).

Han, B. et al. Nano-indentation investigation on the local softening of equiaxed zone in 2060-T8/2099-T83 aluminum-lithium alloys T-joints welded by double-sided laser beam welding. J. Alloy. Compd. 756 , 145–162 (2018).

Li, Z. et al. Bonding and microstructure evolution in electromagnetic pulse welding of hardenable Al alloys. J. Mater. Process. Technol. 290 , 116965 (2021).

Meng, X. M., Bachmann, M., Artinov, A. & Rethmeier, M. Experimental and numerical assessment of weld pool behavior and final microstructure in wire feed laser beam welding with electromagnetic stirring. J. Manufact. Process. 45 , 408–418 (2019).

Fuller, C. B., Mahoney, M. W., Calabrese, M. & Micona, L. Evolution of microstructure and mechanical properties in naturally aged 7050 and 7075 Al friction stir welds. Mater. Sci. Eng. A 9 , 2233–2240 (2010).

Ritchie, R. O. Toughening materials enhancing resistance to fracture. Philos. Trans. R. Soc. A 379 , 20200437 (2021).

Wang, L. et al. A pathway to mitigate macrosegregation of laser-arc hybrid Al-Si welds through beam oscillation. Int. J. Heat. Mass Tran. 151 , 119467 (2020).

Liu, L. et al. Making ultrastrong steel tough by grain-boundary delamination. Science 368 , 1347–1352 (2020).

Shi, P. J. et al. Hierarchical crack buffering triples ductility in eutectic herringbone high-entropy alloys. Science 373 , 912–918 (2020).

Yang, W. J. et al. Achieving high strength and ductility in double-sided friction stir processing 7050-T7451 aluminum alloy. Mater. Sci. Eng. A 707 , 193–198 (2017).

Bocchi, S., D’Urso, G. & Giardini, C. The effect of heat generated on mechanical properties of friction stir welded aluminum alloys. Int. J. Adv. Man. Tech. 112 , 1513–1528 (2021).

Nye, J. F. et al. Some geometrical relations in dislocated cyrstals. Acta Metall. 1 , 153–162 (1953).

Guo, Y. et al. Dislocation density distribution at slip band-grain boundary intersections. Acta Mater. 182 , 172–183 (2020).

Maier, P., Richter, A., Faulkner, R. G. & Ries, R. Application of nanoindentation technique for structural characterisation of weld materials. Mater. Charact. 48 , 329–339 (2002).

Oliver, W. C. & Pharr, G. M. An improved technique for determining hardness and elastic modulus using load and displacement sensing indentation experiments. J. Mater. Res. 7 , 1564–1583 (1992).

Burnett, T. L. et al. Large volume serial sectioning tomography by Xe plasma FIB dual beam microscopy. Ultramicroscopy 161 , 119–129 (2016).

Wang, L. et al. Three-dimensional finite element analysis using crystal plasticity for a parameter study of microstructurally small fatigue crack growth in a AA7075 aluminum alloy. Int. J. Fatigue 31 , 651–658 (2009).

Kords, C. & Raabe, D. On the role of dislocation transport in the constitutive description of crystal plasticity. Berlin: epubli GmbH (2013).

Ardell, A. J. Precipitation hardening. Metall. Trans. A 16 , 2131–2165 (1985).

Hansen, N. Hall-Petch relation and boundary strengthening. Scr. Mater. 51 , 801–806 (2004).

Ma, K. K. et al. Mechanical behavior and strengthening mechanisms in ultrafine grain precipitation-strengthened aluminum alloy. Acta Mater. 62 , 141–155 (2014).

Bailey, J. E. & Hirsch, P. B. The dislocation distribution, flow stress, and stored energy in cold-worked polycrystalline silver. Philos. Mag. 5 , 485–497 (1960).

Download references

Acknowledgements

The authors sincerely thank the National Natural Science Foundation of China (U2032121 and 12192212). As an honorary staff, SC Wu is grateful for his senior visiting position at the Henry Royce Institute for Advanced Materials funded by EPSRC (EP/R00661X). AM Korsunsky thanks the ESPRC support (EP/V007785/1). PJ Withers also acknowledges a European Research Council Advanced Grant (CORREL-CT, Grant No 695638).

Author information

Authors and affiliations.

State Key Laboratory of Traction Power, Southwest Jiaotong University, Chengdu, PR China

Yanan Hu, Shengchuan Wu, Yukuang Yu & Guozheng Kang

School of Mechanics and Aerospace Engineering, Southwest Jiaotong University, Chengdu, PR China

Yanan Hu, Xu Zhang & Guozheng Kang

Henry Royce Institute, Department of Materials, The University of Manchester, Manchester, UK

Shengchuan Wu, Xiangli Zhong & Philip J. Withers

Institute of Metal Research, Chinese Academy of Sciences, Shenyang, PR China

School of Materials Science and Engineering, Shanghai Jiao Tong University, Shanghai, PR China

Zhao Shen & Xiaoqin Zeng

Department of Materials, University of Oxford, Oxford, UK

Zhao Shen & Sergio Lozano-Perez

Department of Engineering Science, University of Oxford, Oxford, UK

Alexander M. Korsunsky

Shanghai Synchrotron Radiation Facility (SSRF), Shanghai Advanced of Sciences, Shanghai, PR China

Yanan Fu & Tiqiao Xiao

Science and Technology on Power Beam Processes Laboratory, AVIC Manufacturing Technology Institute, Beijing, PR China

Zhigang Che

Beijing Synchrotron Radiation Facility (BSRF), Chinese Academy of Sciences, Beijing, PR China

Qingxi Yuan

You can also search for this author in PubMed   Google Scholar

Contributions

Y.N.H., S.C.W., Y.K.Y., T.Q.X., and Q.X.Y. conducted in situ SR-μCT experiments and analyzed the resultant data. Y.N.H., Y.G., S.C.W., Z.S., A.M.K., X.L.Z. and P.J.W. characterized microstructures. Y.N.H., S.C.W., and Y.N.F. performed mechanical tests. X.Z. performed the modeling. Y.N.H., S.C.W. and Z.G.C. were involved with processing development and sample weld. Y.N.H., S.C.W., Z.S., S.L.-P., X.Q.Z., and P.J.W. drafted the initial manuscript. S.C.W., G.Z.K., and P.J.W. conceived, designed, and led the project. All co-authors contributed to the data analyses.

Corresponding authors

Correspondence to Shengchuan Wu , Zhao Shen or Philip J. Withers .

Ethics declarations

Competing interests.

The authors declare no competing interests.

Peer review

Peer review information.

Nature Communications thanks Caiwang Tan and the other, anonymous, reviewers for their contribution to the peer review of this work.  Peer reviewer reports are available.

Additional information

Publisher’s note Springer Nature remains neutral with regard to jurisdictional claims in published maps and institutional affiliations.

Supplementary information

Supplementary information, peer review file, description of additional supplementary files, supplementary movie 1, rights and permissions.

Open Access This article is licensed under a Creative Commons Attribution 4.0 International License, which permits use, sharing, adaptation, distribution and reproduction in any medium or format, as long as you give appropriate credit to the original author(s) and the source, provide a link to the Creative Commons license, and indicate if changes were made. The images or other third party material in this article are included in the article’s Creative Commons license, unless indicated otherwise in a credit line to the material. If material is not included in the article’s Creative Commons license and your intended use is not permitted by statutory regulation or exceeds the permitted use, you will need to obtain permission directly from the copyright holder. To view a copy of this license, visit http://creativecommons.org/licenses/by/4.0/ .

Reprints and permissions

About this article

Cite this article.

Hu, Y., Wu, S., Guo, Y. et al. Inhibiting weld cracking in high-strength aluminium alloys. Nat Commun 13 , 5816 (2022). https://doi.org/10.1038/s41467-022-33188-x

Download citation

Received : 19 April 2022

Accepted : 07 September 2022

Published : 03 October 2022

DOI : https://doi.org/10.1038/s41467-022-33188-x

Share this article

Anyone you share the following link with will be able to read this content:

Sorry, a shareable link is not currently available for this article.

Provided by the Springer Nature SharedIt content-sharing initiative

This article is cited by

Ultrastrong and ductile steel welds achieved by fine interlocking microstructures with film-like retained austenite.

• Joonoh Moon
• Gyuyeol Bae

Nature Communications (2024)

• Shaoning Geng
• Xiangqi Wang

Fatigue Properties of Quenching and Partitioning 980 Steel Laser Welding Joint Obtained under Rapid Cooling Condition

• Jiazhuang Tian
• Shanglu Yang

Journal of Materials Engineering and Performance (2024)

A review of aluminum metal matrix composites: fabrication route, reinforcements, microstructural, mechanical, and corrosion properties

• Ashish Kumar
• Virendra Pratap Singh
• Abdel-Hamid I. Mourad

Journal of Materials Science (2024)

Tensile Deformation Behavior of 6082-T6 Aluminum Alloy After Repair Welding by Metal Inert Gas Process

• Aphichat Sakkaeo
• Suchart Chantaramanee
• Phairote Sungkhaphaitoon

Journal of Materials Engineering and Performance (2023)

By submitting a comment you agree to abide by our Terms and Community Guidelines . If you find something abusive or that does not comply with our terms or guidelines please flag it as inappropriate.

Quick links

• Explore articles by subject
• Guide to authors
• Editorial policies

Sign up for the Nature Briefing newsletter — what matters in science, free to your inbox daily.

• Original Article
• Open access
• Published: 15 October 2022

Numerical Simulation and Experimental Verification of Laser Multi-Section Welding

• Jia Liu   ORCID: orcid.org/0000-0001-7640-2936 1 , 2 , 3 ,
• Tao Jiang 1 , 2 , 3 ,
• Yan Shi 1 , 2 , 3 ,
• Hongyin Zhu 1 , 2 , 3 &
• Yuchi Dai 1 , 2 , 3  

Chinese Journal of Mechanical Engineering volume  35 , Article number:  125 ( 2022 ) Cite this article

2158 Accesses

1 Citations

Metrics details

To address the problems of large welding deformation and splashing in the resistance spot welding of the lubricating oil cooler plate, the laser spot welding was employed, instead of the resistance spot welding, and a novel laser spot welding was proposed, i.e., laser multi-section welding. The major processes involved in this study referred to a finite element model of pulsed laser lap welding built by adopting SYSWELD simulation software, as well as the laser welding of various welding methods. The effect of different welding methods on the welding quality was studied, the parameters of the average power and the duty cycle were optimized in line with the comparative analysis of the experimentally achieved results and the numerical simulation. As indicated from the experimentally achieved results, when the new 6-sections welding method was adopted, the resulting welded joint achieved the most uniform heat input and the largest welding fusion area, and the tensile properties exhibited by the welded joints were significantly enhanced, whereas some pores remained. By altering the duty cycle, pores could be eliminated to further improve the quality of the joint. The mentioned process method could tackle down the problems facing conventional resistance spot welding. Furthermore, it was capable of improving the uneven heat input of the laser spot welding.

1 Introduction

A lubricating oil cooler refers to an extensively applied heat exchanger that critically impacts nuclear power, chemical industry and other relevant fields, which can reduce energy loss to a certain extent [ 1 ]. Figure 1 presents a structural diagram of a lubricating oil cooler. Austenitic stainless steel has been broadly applied in railway vehicles and the heat exchanger/chemical industry for their prominent corrosion resistance, high strength and efficient stamping performance [ 2 , 3 , 4 , 5 ]. Currently, the lube oil cooler plate primarily employs the resistance spot welding for connection, whereas serious spatter issues exist in such a welding process, As shown in Figure 2 . Besides, the resistance spot welding exhibits the defects below. First, when welding one of the points, a part of the current tends to flow through the welded solder joints, which reduces the quality of the welding. Next, the resistance of the electrode and the workpiece are easy to be affected by considerable factors (e.g. such as temperature, electrode pressure and other material properties), thereby causing the resistance spot welding to be extremely vulnerable to outside conditions, Moreover, the requirements of cleanliness are high. In the presence of impurities on the surface, the resistance between the weldments will increase significantly, thereby directly affecting its conductivity features. For the mentioned reasons, the laser welding technology is proposed to replace the original resistance spot welding.

Lubricating-oil cooler

Resistance spot welding process and internal section photos of the sealed cavity: a Spatter during welding; b Deposition of splashes in the cavity

Laser welding is recognized as an efficient and precise welding method that exploits a high energy density laser beam as the heat source. It is advantageous for its high energy density, as well as for its precise focus, strong penetrating ability, high welding efficiency and applicability since applies to a myriad of materials. Currently, it is mostly employed in aerospace, automotive, microelectronics and nuclear industries with high precision and high-quality welding areas [ 6 , 7 , 8 , 9 ]. Compared with resistance spot welding, laser spot welding has many advantages, such as high welding accuracy, accurate energy control, and good penetration. With the use of the non-contact welding, the welding quality can be ensured for the welding of complex structures, and the application turns out to be more flexible. However, the laser spot welding, as a cutting-edge spot welding method, still faces many unsolved technical problems. The laser spot welding will cause stress concentration and uneven heat inputs. Moreover, as impacted by the extremely fast heating and cooling speeds, extreme defects (e.g., pores and cracks) are easy to produce, which seriously reduces the quality of the weld formation and the weld strength [ 10 , 11 , 12 , 13 ].

Finite element simulation can be adopted at the initial stage of design since it acts as a powerful tool for experiment and optimization in virtual environments. Besides, such a technique can help reduce the number of experiments and lay a theoretical basis for strength analyses [ 14 , 15 , 16 , 17 ].

Yan et al. [ 18 ] studied the structure and mechanical properties of 304 stainless steel joints welded by tungsten inert gas (TIG) welding, laser welding, and laser-TIG hybrid welding. The results of his experimentations show that the laser-welded joint has the highest tensile strength and the smallest dendrite size among all, while the TIG welded joint has the lowest tensile strength and the largest dendrite size. Chakravarthy et al. [ 19 ] also studied the effect of laser welding process parameters and welding speed on the welding performance of 70/30 Cu-Ni alloy welded joints. His processes proved that, when the laser uses a welding speed of 1.5 m/min, the joints produced fine, equiaxed, and uniformly distributed grains that have better mechanical properties than other joints in the fusion zone. Going further, Bowen Liu et al. [ 20 ] studied the dual laser beam welding of 316L austenitic stainless steel, using an artificial neural network (ANN) and genetic algorithm (GA) to obtain the best welding parameters of them. Some other studies have shown that porosity features are significantly reduced compared to the original weld. Derakhshan et al. [ 21 ] examined the effects of heat input on mechanical properties obtained by four different welding processes: Autogenous Laser Welding, Cold Wire Assisted Laser Welding, Hybrid Laser Arc Welding, and traditional Submerged Arc Welding. SYSWELD software was used to predict the temperature field, residual stress, and deformation and the experimental results had a satisfactory correlation with the calculated results. The process proved that, by increasing the cooling rate of the welding process, it is possible to effectively improve the microhardness of the weld center. Adding to that, Saravanan et al. [ 22 ] compared the macrostructures of Hastelloy C-276 joints obtained by Nd: YAG laser butt-welding through numerical simulation and experimental methods. The results show that SYSWELD is suitable for simulating the weld profile and temperature distribution of laser-welded Hastelloy C-276 joints. In this case, the profile obtained by the numerical simulation was also very consistent with the experimental results. Last but not least, Ganesh et al. [ 23 ] performed the thermo-mechanic analysis using SYSWELD and both residual stress and distortion predicted by FEM simulation were in acceptable agreement with the experimental measurements.

Since the lubricating oil cooler has higher requirements for heat exchange efficiency, it is desirable to have the greatest strength under the identical welding area to avoid the excessive welding area from affecting the flow of the fluid media and affecting the heat transfer. On that basis, this study proposed a novel laser welding method (multi-section welding) for the plate heat exchanger welding. Such a welding method was used instead of the laser spot welding. The welding path was segmented, and its laser parameters were optimized, which avoided the spatter produced by the conventional spot welding. Moreover, the welding heat input turned out to be uniform and satisfied the requirements of use. This study primarily employed a combination of numerical simulation and experiments to study the effect of different welding methods on the weld pool morphology, the fracture morphology and the mechanical properties of the joint. Next, the parameters were further optimized. Hopefully, the study of such a laser welding method can provide researchers and industrial engineers with more insights.

2 Materials and Methods

First, 304 austenitic stainless steel plates with the dimension of 100 mm×15 mm×2 mm acted as the welding base metal. The length of the lap joint area was 35 mm, and the diameter of the circular weld was 4 mm, as shown in Figure 3 . The detailed composition of base metal is listed in Table 1 . Before the welding, the impurities on the surface of the workpiece should be cleaned with acetone reagent.

Geometric model

The laser welding system was composed of TruDisk 8002 laser and KUKA robot. After the welding, the sample was cut to perform a tensile test. Besides, there are three samples for the respective experimental parameter. The average value was taken to determine the final average tensile shear force. Furthermore, a stereomicroscope was employed to photograph the fracture morphology of the stretched sample and to observe the morphology and fusion state of the combined surface of the welded joint. Moreover, a metallographic sample was prepared, and a metallographic microscope was employed to characterize the weld morphology under a metallographic microscope, as well as to measure the depth and width of fusion simultaneously. The observation method is illustrated in Figure 4 . Such an observation direction was applied when different laser welding methods were employed.

Schematic diagram of the observation direction of the metallographic sample

To ensure the weld strength, the overall penetration depth of the laser welding was controlled in the range of 3±0.2 mm. The different power parameters for segmental welding are listed in Table 2 . The welding sequence from the center of the circle to the end of it and four welding methods (i.e., 1-section, 4-sections, 5-sections, and 6-sections) were adopted. The detailed schematic diagram of welding is shown in Figure 5 .

Schematic diagram of laser overlap welding method: a 1-section; b 2-sections; c 3-sections; d 4-sections

3 Finite Element Analysis

3.1 establishment of heat source modes.

Notably, the heat source should be checked prior to the simulation calculation to make the result more accurate. The 3D Gaussian heat source was selected regardless of the flow of the molten pool, and the energy distribution of the temperature field of the source is presented in Figure 6 . According to the figure, the heat flow energy in the middle was high, and the surroundings were low, so the heat source might display a Gaussian distribution.

3D Gaussian heat source

In Figure 6 , \({Q}_{0}\) denotes the maximum input energy density; \({r}_{e}\) and \({r}_{i}\) are Gaussian parameters, where \({r}_{e}\) represents the maximum characteristic radius on the plane \(z\)  =  \({z}_{e}\) , and \({r}_{i}\) denotes the maximum characteristic radius on the plane \(z\)  =  \({z}_{i}\) ;  \({z}_{e}\) and \({z}_{i}\) express are the 3D Gaussian heat sources. The position makes a parameter with the surface of the piece to be welded, where \({z}_{i}\) denotes the relative height difference between the upper surface of the 3D Gauss heat source and the surface of the workpiece, and \({z}_{i}\) is the relative height difference between the lower surface of the 3D Gauss and the surface, positive above the outward. This is also likely to be negative. Thus, the parameter relation can be expressed as [ 24 , 25 , 26 , 27 ]:

For the numerical simulation of the laser welding, the total energy of the 3D Gaussian heat source is written as:

Therefore, when using a 3D Gaussian heat source to simulate laser welding, the energy distribution is represented as:

In Eq. ( 1 ), \({r}_{0}\left(z\right)\) is the maximum radius of action on the plane of height \(z\) . Eq. ( 2 ) can be used to calculate the total energy of the heat source when using a 3D Gaussian heat source to simulate laser welding. In this case, Q is the total energy of the heat source; \(\eta\) is the absorption rate of the metal to the laser and P stands for the laser power. Last, Eq. ( 3 ) addresses the energy distribution of 3D Gaussian heat source simulating laser welding where \(q\left(r,z\right)\) represents the energy density from the center r on the plane of height \(z\) and H is the total height of the heat source.

In SYSWELD, the heat source parameters can be calculated from the actual weld profile. Here, the welded joint profile parameters of the single-pass weld optimized by previous experiments are used as the basis for calculating the heat source parameters, which can effectively reduce the optimization of the heat source parameters. The morphology of the single-pass weld pool was obtained with a laser power of 2500 W, a pulse frequency of 40 Hz, and a welding speed of 1.5 m/min is shown in Figure 7 . The weld width is 2.35 mm, the penetration depth is 2.94 mm, and the joint surface width is 1.2 mm. Input data to SYSWELD Among them, the heat source parameters obtained by the feedback were \({r}_{e}\)  = 1.175, \({r}_{i}\)  = 0.6, \({z}_{e}\)  = 0, \({z}_{i}\)  = − 2.94. This heat source parameter was applied to laser multi-section welding. Using the 1-section welding method, the laser parameters included the laser power of 1800 W, the defocus of + 4 mm, the welding speed of 1.5 m/min. According to Figure 8 , as indicated from the comparison, the morphology of the molten pool in the numerical simulation was identical to that of the actual welded joint, the accuracy of the selected heat source model and heat source parameters could be verified.

Shape of the welded joint of single-pass weld

Comparison of weld morphology of welded joints: a Numerical simulation of the molten pool morphology; b Actual weld joint morphology

3.2 Establishment of the Finite Element Model

The energy density generated by the laser welding is extremely high, and the heating speed is fast. When the weld is locally heated, the temperature difference between the weld and the surrounding area turns out to be significantly large. However, in the area far away from the weld, the temperature distribution is basically insignificant. Based on the characteristics of the temperature distribution of the laser welding, when meshing, a finer mesh is required in the weld area and its vicinity. The meshing result of the welded specimen is presented in Figure 9 . To improve the calculation accuracy, the grid division should be relatively fine in and fit around the weld zone. Here, the minimum grid size was 0.2 mm during the grid division, so the result of the temperature field was more accurate. In the area far from the weld, the maximum mesh size reached 4 mm, thereby leading to a reduced number of grids and calculation time, while the small and big mesh were connected via the transition mesh.

Meshing result of the welded specimen: a Grid within the weld; b Overall grid

4 Results and Discussion

4.1 numerical simulation and experimental comparison.

The welding experiments were performed according to the welding parameters (Table 2 ). In addition, the simulated molten pool morphology should be compared with the actual joint morphology when the welding experiment was completed. The diameter of the circular weld was 4 mm, so a cylinder with a diameter of 4 mm was taken for the temperature field observation, and the change depth of fusion could be observed directly. The temperature field cloud diagram obtained when the welding method was 1-section is presented in Figure 10 . According to the figure, the depth of fusion tended to increase, with a change range of 3‒3.38 mm. The alterations were attributed to the different temperatures on the workpiece, as impacted by the action of the moving heat source during the welding. The moving heat source exhibited a significant performance on each position of the welding due to the circular welding path and the small diameter of the circumference. Accordingly, the secondary heating exerted a superimposed effect on the weld, thereby resulting in a greater change in the depth of fusion.

Temperature field distribution cloud map under the welding condition of 1-section

The joint morphology with the welding mode of 1-section shown in Figure 11 helped to closely observe the differences in the depth of fusion. In the figure, the molten pool morphology of numerical simulation was identical to the morphology of the actual welded joint, but the range of weld penetration is way too large to meet the experimental requirements. Besides, no further discussion would be conducted.

Welding method is a 1-section of joint morphology: a Numerical simulation of molten pool morphology; b Actual welded joint morphology

Considering the 4, 5, and 6-sections of welding, the temperature field distribution cloud diagrams obtained are shown in Figures 12 , 13 and 14 , respectively. As observed from the figures, the welding process was the most stable one and the variation range (3‒3.12 mm) of the depth of fusion was the smallest when the 6-section welding method was adopted. In addition, the morphology of the welded joint is presented in Figures 15 , 16 and 17 , thereby proving that the depth of fusion was maintained at about 3 mm. The optimal welding effect was achieved when the 6-sections welding method was adopted.

Temperature field distribution cloud map under the welding condition of 4-sections

Temperature field distribution cloud map under the welding condition of 5-sections

Temperature field distribution cloud map under the welding condition of 6-sections

Welding method is 4-sections of joint morphology: a Numerical simulation of molten pool morphology; b Actual welded joint morphology

Welding method is 5-sections of joint morphology: a Numerical simulation of molten pool morphology; b Actual welded joint morphology

Welding method is 6-sections of joint morphology: a Numerical simulation of molten pool morphology; b Actual welded joint morphology

Figure 18 illustrates the temperature field of the welded combined surface and in the area enclosed by the welding path, the overall fusion was better. However, there is still a small part of the unfused areas. It was also observed that the incomplete fused positions at the combined surface of the welding joint were all around the welding starting point. The welding method of the 6-sections changed to a new welding path as shown in Figure 19 , that is, the length of the initial welding section is increased from 2 mm to 3 mm (new 6-sections), while the other paths remained the same. The corresponding average power of each section also changed—to 2500/2350/2150/1975/1600/725 W–while other process parameters had no alterations. The final welding joint morphology obtained is shown in Figure 20 , and it was achieved after a numerical simulation calculation of the temperature field and laser welding experiment. In this stage, the morphological properties are consistent with those of the actual welding joint, and the depth of fusion of the molten pool is about 3 mm. Figure 21 is the combined surface temperature field of the welded joint obtained by the new 6-sections welding method, the welded combined surface was completely fused. which might meet all the experimental requirements.

Temperature field of combined surface under different welding methods: a 4-sections; b 5-sections; c 6-sections

Welding method is the new 6-sections welding path

Weld joint morphology of new 6-sections: a Numerical simulation of molten pool morphology; b Actual welded joint morphology

Welding method is the combined surface temperature field of the new 6-sections welding joint

4.2 Mechanical Properties

To study the effect of different welding methods on the mechanical properties of weld joints, tensile experiments were performed on welded joints obtained under sections 4, 5, 6, and new 6-sections. According to Figure 22 , the tensile shear force obtained by the 6-section welding method was significantly improved compared with the 4-sections and 5-sections welding methods, which reached over 10256 N. Under the new 6-sections welding method, the tensile shear force performance of the welded joint was further improved, the average tensile shear force of the welded joint was 10454 N, and the maximum tensile force increased by 2% compared with the one obtained by 6-section welding. It was therefore indicated that the mechanical properties of this process were up to the mark.

The maximum tensile force of welded joints under different welding methods

To explore the reasons for the different tensile properties caused by using different welding methods, it was necessary to conduct a numerical simulation analysis and an experimental verification of the temperature field at the joint surfaces should be performed. The way of tensile fracture is shown in Figure 23 , fracture positions were all at the combined surface. Figure 24 shows the tensile fracture morphology of the welded joint, and the existence of the area without fusion on the combined surface of the weld was consistent with the results of the temperature field simulation in Figures 18 and 21 . Also, the area of the fusion zone obtained in the 4, 5, 6, and new 6-sections methods were 15.95 mm 2 , 15.88 mm 2 , 17.50 mm 2 , and 17.97 mm 2 , respectively, as shown in Figure 25 . Thus, it can be concluded that the fusion area of combined surface and mechanical properties of welded joints was greatly improved under the welding method based on the new 6-sections welding.

Schematic diagram of tensile specimen fracture

Tensile fracture morphology of welded joint under different welding methods: a 4-sections; b 5-sections; c 6-sections; d new 6-sections

Fusion area of the combined surface of the welded joint under different welding methods

4.3 Process Parameter Optimization

After the observation and analysis of the welded joint obtained by using the new 6-sections welding method, a small number of pores could still be identified. In the pulse laser welding, laser transmission behavior showed significant differences with the variation of the duty cycle, which significantly affected the weld appearance and formation of stomata [ 28 , 29 , 30 ]. The porosity was reduced by changing the duty cycle, thereby further improving the quality of the welded joint. The specific welding process parameters are listed in Table 3 .

Figure 26 presents the weld morphology of welded joints under different duty cycles. Figures 27 and 28 , respectively represent the fusion area and tensile shear force of the welded joint also in different duty cycles. According to the figures, when the duty cycle is increased from 35% to 46%, the fusion area of the welded joint remained basically the same—maintained at about 18.5 mm 2 . However, when the duty cycle was 35%, the tensile shear of the welded joint was significantly reduced. When the duty cycle was small, a high-power density laser beam acted on the workpiece to instantly vaporize the metal to form a keyhole, while considerable plasma was generated to shield the absorption of laser energy by the molten pool with the increase in the duty cycle. Though the power density of the laser was weakened, the plasma generated was also reduced, whereas the absorption rate of the laser energy by the molten pool increased. Lastly, the heat obtained at the combined surface was not so divergent, thereby causing the insignificant change of the fusion area at the combined surface. When the duty cycle was 35%, the power density of the laser spot was extremely high, and a large amount of metal vapor was generated. Due to the small pulse width, the laser acted on the workpiece only for a short time, and the weld pool rapidly cooled and solidified. At this stage, it was too late for the metal vapor to escape, making it easier to form porosity defects. Combined with the tensile fracture morphology of the welded joint shown in Figure 29 , it can be said that there were larger porosity defects at the tensile fracture with a duty cycle of 35%, therefore, the tensile properties of the welded joint were reduced.

Weld morphology of welded joints under different duty cycles: a 35%; b 42%; c 46%; d 50%; e 54%; f 58%; g 75%

Tensile shear force of welded joints under different duty cycles

Fusion area of the combined surface of the welded joint under different duty cycles

Tensile fracture morphology of welded joints under different duty cycles: a 35%; b 42%; c 46%; d 50%; e 54%; f 58%; g 75%

During the duty cycle from 46% to 75%, as it increased, the tensile shear force of the welded joint was gradually reduced. An increased duty cycle caused the laser power density to continue to downgrade, thereby resulting in a reduction in the fusion area of the combined surface. When the duty cycle was between 54% and 58%, the fusion area was nearly equal, whereas the tensile shear force of the welded joint differed by about 220 N. This was because in this scenario, the power density of the pulsed laser was lower, so the generated energy was insufficient to maintain the stable existence of the keyhole. Consequently, the keyhole collapse was easy to occur, thereby causing the gas to stay in the weld, which led to the formation of the pore defect. This phenomenon happens precisely due to the pores at the tensile fracture that the welded joint was pulled, causing the reduction of tensile shear force. When the duty cycle is increased to 75%, the laser power density is extremely low. It can be seen from the topography that there are large unfused areas and multiple pore defects at the joint surface, thereby weakening the tensile properties of welded joints. Based on the above analysis it is possible to state that, when the selected duty cycle is 42% to 46%, the best quality welded joint is obtained with a maximum tensile shear force of about 10632 N. As shown in Figure 30 , the new 6-sections welding method and the 45% duty cycle were adopted to conduct the welding experiments on the trial plate pairs of the lubricating oil cooler. The welding seam quality was good, and there was no splash in the welding cavity, which fully met the experiment.

Welding seam morphology of the plate-pair laser welding of the lubricating oil cooler: a The front side of the weld; b The back of the weld

5 Conclusions

Given the major results of the experiments conducted here, some main conclusions could be drawn after all the processes were achieved:

By comparing the simulation results of different welding methods, it was concluded that the optimal welding was the new 6-sections welding, which exhibited the smallest fluctuation range of depth of fusion. Its experimental results were well consistent with the simulation results.

By performing the tensile test, the mechanical properties of the welded joints under different methods were obtained. With the use of the new 6-sections method, the tensile shear force was the largest. Moreover, the temperature field of the combined surface of the welded joint and the macroscopic morphology of the actual tensile fracture was analyzed. When the welding method was the new 6-section, there was no unmelted area on the combined surface of the welded joint, which also complied with the temperature field results of the numerical simulation.

The variation in the duty cycle affected the pulse width of the laser. When the duty cycle ranged from 42% to 46%, the laser power density was high, the fusion area of the combined surface was large, and the formed keyhole was stable, so it hindered the formation of porosity. In this case, the resulting welded joints exhibited the highest quality.

M Shirzad, M A Delavar, S S M Ajarostaghi, et al. Evaluation the effects of geometrical parameters on the performance of pillow plate heat exchanger. Chemical Engineering Research and Design , 2019, 150: 74-83.

Article   Google Scholar  

H Kumar, N K Singh. Performance of activated TIG welding in 304 austenitic stainless steel welds.  Materials Today Proceedings , 2017, 4(9): 9914-9918.

S S Kumar, N Murugan, K K Ramachandran. Microstructure and mechanical properties of friction stir welded AISI 316L austenitic stainless steel joints . Journal of Materials Processing Technology , 2018, 254: 79-90.

S Singh, K Hurtig, J Andersson. Investigation on effect of welding parameters on solidification cracking of austenitic stainless steel 314. Procedia Manufacturing , 2018, 25: 351-357.

M Safari, H Mostaan, A Ghaderi. Dissimilar resistance spot welding of AISI 304 to AISI 409 stainless steels: Mechanical properties and microstructural evolutions. Metallurgical Research and Technology , 2018, 115(6): 610.

Q Wu, J Gong, G Chen, et al. Research on laser welding of vehicle body. Optics & Laser Technology , 2008, 40(2): 420-426.

M Gao, Y Kawahito, S Kajii. Observation and understanding in laser welding of pure titanium at subatmospheric pressure. Optics Express , 2017, 25(12): 13539–13548.

I Miyamoto, Y Okamoto, A Hansen, et al. High speed, high strength microwelding of Si/glass using ps-laser pulses. Optics express , 2015, 23(3): 3427-3439.

Z Y Huang, Z Luo, S S Ao, et al. Effect of Laser Welding Parameters on Weld Bowing Distortion of Thin Plates. High Temperature Materials and Processes , 2018, 37(4): 299-311.

M Masoumi, S P H Marashi, M Pouranvari. Metallurgical and Mechanical Characterization of Laser Spot Welded Low Carbon Steel Sheets. Steel Research International , 2010, 81(12): 1144-1150.

Y S Yang, S H Lee. A study on the joining strength of laser spot welding for automotive applications. Journal of Materials Processing Technology , 1999, 94(2-3): 151-156.

B J Aalderink, B Pathiraj, R G K M Aarts. Seam gap bridging of laser based processes for the welding of aluminium sheets for industrial applications. The International Journal of Advanced Manufacturing Technology , 2010, 48(1): 143-154.

J Kundu, T Ray, A Kundu, et al. Effect of the laser power on the mechanical performance of the laser spot welds in dual phase steels. Journal of Materials Processing Technology , 2018, 267.

L F Zhao, C D Shao, Y Takashima, et al. Numerical investigation on fracture initiation properties of interface crack in dissimilar steel welded joints. Chinese Journal of Mechanical Engineering , 2020, 33: 27.

M Dal, R Fabbro. An overview of the state of art in laser welding simulation. Optics & Laser Technology , 2016, 78: 2-14.

K R Balasubramanian, T Suthakar, K Sankaranarayanasamy, et al. Laser Welding Simulations of Stainless Steel Joints Using Finite Element Analysis. Advanced Materials Research , 2012, 383-390: 6225-6230.

Google Scholar  

D W Zhao, D X Ren, K M Zhao, et al. Ultrasonic welding of magnesium–titanium dissimilar metals: A study on thermo-mechanical analyses of welding process by experimentation and finite element method. Chinese Journal of Mechanical Engineering , 2019, 32: 97.

J Yan, M Gao, X Y Zeng. Study on microstructure and mechanical properties of 304 stainless steel joints by TIG, laser and laser-TIG hybrid welding. Optics and Lasers in Engineering , 2010, 48(4): 512-517.

M P Chakravarthy, N Ramanaiah, B S S Rao. Effect of laser welding on mechanical properties of 70/30 Cu-Ni alloy welds. Proceedings of the Institution of Mechanical Engineers, Part B: Journal of Engineering Manufacture , 2014, 228(9): 1153-1161.

B W Liu, W T Jin, A J Lu, et al. Optimal design for dual laser beam butt welding process parameter using artificial neural networks and genetic algorithm for SUS316L austenitic stainless steel. Optics & Laser Technology , 2020, 125: 106027.

E D Derakhshan, N Yazdian, B Craft, et al. Numerical simulation and experimental validation of residual stress and welding distortion induced by laser-based welding processes of thin structural steel plates in butt joint configuration. Optics & Laser Technology , 2018, 104: 170–182.

S Saravanan, K Raghukandan, G S Kumaret. Comparison of numerical and experimental macrostructure in Nd: YAG laser welding of Hastelloy C-276. Optik, 2019, 180: 562-568.

K C Ganesh, M Vasudevan, K R Balasubramanian, et al. Modeling, prediction and validation of thermal cycles, residual stress and distortion in type 316 LN stainless steel weld joint made by TIG welding process. Procedia Engineering , 2014, 86: 767-774.

N Astafeva. Modeling of welding process and allied technologies. Key Engineering Materials , 2018, 69: 323-328.

T Kik, J Górka. Numerical simulations of laser and hybrid S700MC T-joint welding. Materials , 2019, 12(3): 516.

C S Wu, H G Wang, Y M Zhang. A new heat source model for keyhole plasma arc welding in FEM analysis of the temperature profile. Welding Journal , 2006, 85(12): 284-291.

T Kik. Computational techniques in numerical simulations of arc and laser welding processes. Materials , 2020, 13(3): 608.

W Meng, Z L Xu, Q S Ma, et al. Pulse fiber laser welding of AISI 321-AISI 405 stainless steel thick plates butt joints. Journal of Materials Processing Technology , 2019, 271: 214-225.

Z Yang, B J Qi, B Q Cong, et al. Microstructure, tensile properties of Ti-6Al-4V by ultra high pulse frequency GTAW with low duty cycle. Journal of Materials Processing Technology , 2015, 216: 37-47.

K Y Wu, Z Y Liang, T Yin, et al. Double pulse low-frequency modulation for high-power double-wire pulsed GMAW. Journal of Manufacturing Science and Engineering , 2018, 140(9): 091004.

Download references

Acknowledgments

Not applicable.

Supported by Department of Science and Technology of Jilin Province of China (Grant No. 20180201063GX).

Author information

Authors and affiliations.

School of Electromechanical Engineering, Changchun University of Science and Technology, Changchun, 130022, China

Jia Liu, Tao Jiang, Yan Shi, Hongyin Zhu & Yuchi Dai

Engineering Research Center of Laser Processing for Universities of Jilin Province, Changchun, 130022, China

National Base of International Science and Technology Cooperation in Optics, Changchun, 130022, China

You can also search for this author in PubMed   Google Scholar

Contributions

JL was in charge of the whole trial; TJ wrote the manuscript; HYZ and YCD assisted with sampling and laboratory analyses. YS supervised the study. All authors read and approved the final manuscript.

Authors’ Information

Jia Liu, born in 1983, is currently an associate professor at Changchun University of Science and Technology, China . He received his PhD degree from Changchun University of Science and Technology, China , in 2012. His research interests include laser welding, laser composite processing, and its finite element simulation technology.

Tao Jiang, born in 1992, is currently a master candidate at Engineering Research Center of Laser Processing , Changchun University of Science and Technology, China .

Yan Shi, born in 1972, is currently a professor at Changchun University of Science and Technology, China . He received his PhD degree from Changchun University of Science and Technology, China , in 2007.

Hongyin Zhu, born in 1994, is currently a master candidate at Engineering Research Center of Laser Processing , Changchun University of Science and Technology, China .

Yuchi Dai, born in 1994, is currently a PhD candidate at Engineering Research Center of Laser Processing, Changchun University of Science and Technology, China .

Corresponding author

Correspondence to Jia Liu .

Ethics declarations

Competing interests.

The authors declare no competing financial interests.

Rights and permissions

Open Access This article is licensed under a Creative Commons Attribution 4.0 International License, which permits use, sharing, adaptation, distribution and reproduction in any medium or format, as long as you give appropriate credit to the original author(s) and the source, provide a link to the Creative Commons licence, and indicate if changes were made. The images or other third party material in this article are included in the article's Creative Commons licence, unless indicated otherwise in a credit line to the material. If material is not included in the article's Creative Commons licence and your intended use is not permitted by statutory regulation or exceeds the permitted use, you will need to obtain permission directly from the copyright holder. To view a copy of this licence, visit http://creativecommons.org/licenses/by/4.0/ .

Reprints and permissions

About this article

Cite this article.

Liu, J., Jiang, T., Shi, Y. et al. Numerical Simulation and Experimental Verification of Laser Multi-Section Welding. Chin. J. Mech. Eng. 35 , 125 (2022). https://doi.org/10.1186/s10033-022-00797-y

Download citation

Received : 22 November 2020

Revised : 20 May 2021

Accepted : 22 April 2022

Published : 15 October 2022

DOI : https://doi.org/10.1186/s10033-022-00797-y

Share this article

Anyone you share the following link with will be able to read this content:

Sorry, a shareable link is not currently available for this article.

Provided by the Springer Nature SharedIt content-sharing initiative

• Laser multi-section welding
• Finite element analysis
• Mechanical properties

• [email protected]
• ( +1) 832-234-9188
• 2220 O'Toole Ave, San Jose, CA 95131, USA
• 4224 Clay Business Dr, Katy, TX 77449

Glass Laser Welding Process Experiment for OLED Packaging

• August 15, 2023

Introduction

As the demand for cutting-edge display technologies escalates, the pursuit of efficient and reliable packaging methods for Organic Light Emitting Displays (OLEDs) intensifies. In this quest, the glass laser welding process emerges as a technological breakthrough that holds the potential to redefine OLED packaging. In this technical article, we delve into a comprehensive experiment that harnesses ultrafast lasers to perform glass laser welding for OLED packaging. This experimental endeavor not only showcases the prowess of femtosecond-level ultrafast lasers but also sheds light on the groundbreaking Han’s Laser glass ceramic fusion welding processing system, poised to transform the flat panel display industry.

Principle of welding

Experiment Parameters and Setup

The experiment centers on a meticulous process that demonstrates the capabilities of the glass laser welding Process. A 0.3mm thick glass plate, with a 20mm×20mm welding area. The employed 20W femtosecond laser is integrated into the independently developed Han’s Laser welding machine . This machine boasts a marble movement platform that ensures pinpoint accuracy, a pivotal factor in achieving flawless welds.

OLED display basic structure

The Sequential Steps

Pristine Surface Preparation: The experiment commences with a meticulous cleaning process. Employing an ultrasonic cleaning system, the glass surface undergoes thorough cleansing, followed by a sequence of alcohol and distilled water wipes. The surface is then meticulously dried, setting the stage for optimal welding conditions.

Glass Lamination: The glass sheets are securely positioned using a specialized fixture. Precise pressure adjustment ensures that the distance between the two substrate glasses remains below a quarter of the laser wavelength. This meticulous optical bonding guarantees the integrity of the ensuing weld.

Focus and Alignment: The marble motion platform comes into play as the Z-axis height is adjusted to achieve laser focus. The laser welding head’s alignment is carefully calibrated, ensuring its focal point resides at the glass frit between the two substrates, as predetermined by the laser working distance.

Laser Precision: With the apparatus meticulously prepared, the experiment proceeds to the laser welding phase. The software-driven process leverages adjustable process parameters, encompassing laser attributes, welding speed, trajectory, and more. This orchestrated symphony culminates in the precise welding of the glass along the predefined trajectory.

Weld Seam Morphology

Ultrafast Laser Advantages

Central to the experiment’s success is the utilization of femtosecond-level ultrafast lasers. These lasers, boasting pulse durations in the order of picoseconds and femtoseconds, exemplify high peak energy and minimal thermal effects. This combination empowers direct welding between glass components, eliminating the need for glass frit screen printing and intricate heat dissipation structures. The result is a fusion of glass that epitomizes uniformity, void of defects such as pores and stress deformations. Furthermore, the thermal impact is negligible, safeguarding the delicate components housed within the glass and adhering to stringent airtightness prerequisites essential for OLED packaging.

Welding surface

Han’s Laser Glass Ceramic Fusion Welding Processing System

The highlight of this journey is the Han’s Laser glass ceramic fusion welding processing system. Anchored by an ultra-fast laser (femtosecond/picosecond) and a high-precision motion control system, this innovation finds its footing in OLED screen packaging, precision welding, and glass-metal fusion. With precision, control, and minimal thermal impact, this system opens vistas of possibility within the domain of transparent and brittle material welding.

The glass laser welding process experiment presents a paradigm shift in OLED packaging, exemplifying the potency of ultrafast lasers and innovative welding methodologies. As the Han’s Laser glass ceramic fusion welding processing system heralds a new era of precision welding, the future of OLED packaging stands at the threshold of unprecedented reliability and efficiency. This endeavor not only contributes to the evolution of display technology but also signifies the continuous pursuit of excellence in precision engineering.

Error: Contact form not found.

share this!

August 7, 2024

This article has been reviewed according to Science X's editorial process and policies . Editors have highlighted the following attributes while ensuring the content's credibility:

fact-checked

peer-reviewed publication

trusted source

Lasers deliver powerful shocking punch in material experiments

by Anne M. Stark, Lawrence Livermore National Laboratory

Shock experiments are widely used to understand the mechanical and electronic properties of matter under extreme conditions, like planetary impacts by meteorites. However, after the shock occurs, a clear description of the post-shock thermal state and its impacts on material properties is still lacking.

Lawrence Livermore National Laboratory (LLNL) scientists used ultra-fast X-ray probes to track the thermal response of aluminum and zirconium on shock release from experiments and found the resulting temperatures were much higher than expected. The research is published in the Journal of Applied Physics .

A shock wave is a large-amplitude mechanical wave across which pressure, density, particle velocity, temperature and other material properties change abruptly as the wave travels through the material. The shock compression process is thermodynamically irreversible, where a substantial portion of the energy in a shock wave goes into raising the entropy and temperature of the material.

The team used diffraction patterns from 100-femtosecond X-ray pulses to investigate the temperature evolution of laser-shocked aluminum-zirconium metal film composites at time delays ranging from 5 to 75 nanoseconds driven by a 120-picosecond short-pulse laser.

"We found significant heating of both the aluminum and zirconium after shock release, which can be attributed to heat generated by inelastic deformation," said LLNL principal investigator Harry Radousky, a co-author of the study.

As it turns out, a conventional hydrodynamic model that uses typical descriptions of aluminum and zirconium mechanical strength and elevated strength responses (which might be attributed to an unknown strain rate) did not fully account for the measured temperature increase, which suggests that other strength-related mechanisms could play an important role in thermal responses under shock wave loading/unloading cycles.

"What we found is that a significant portion of total shock energy delivered by lasers become heat due to defect-facilitated plastic work, leaving less converted to kinetic energy ," said LLNL scientist Mike Armstrong, another co-author of the study. "This heating effect may be common in laser-shocked experiments but has not been well acknowledged. The high post-shock temperatures may induce phase transformation of materials during shock release."

Armstrong said another potential application of the study is preserving magnetic records from planetary surfaces that have a shock history from frequent impact events.

Using the Matter in Extreme Conditions instrument at the Linac Coherent Light Source, the team found results showed much higher residual temperatures than expected from standard hydrodynamic release simulations, indicating there are other heat generating processes (void formation is one example) occurring during the release not usually included in these models.

Journal information: Journal of Applied Physics

Provided by Lawrence Livermore National Laboratory

Explore further

Feedback to editors

Flood of 'junk': How AI is changing scientific publishing

3 hours ago

135-million-year-old marine crocodile sheds light on Cretaceous life

13 hours ago

Researchers discover new material for optically-controlled magnetic memory

A new mechanism for shaping animal tissues

15 hours ago

NASA tests deployment of Roman Space Telescope's 'visor'

16 hours ago

How do butterflies stick to branches during metamorphosis?

17 hours ago

Historic fires trapped in Antarctic ice yield key information for climate models

Hubble spotlights a supernova

18 hours ago

New technology uses light to engrave erasable 3D images

Seagrasses filter human pathogens in marine waters

19 hours ago

Relevant PhysicsForums posts

The double-slit experiment with a pit in the screen.

Aug 8, 2024

How to observe difference of L/R rotation intensity of CPL

Jul 29, 2024

Increase in numerical aperture leads to a decrease in line width?

Jul 19, 2024

Ray Tracing Software for Broadband Lasers

Jul 18, 2024

The detection of the carrier-envelope offset frequency (fCEO) of optical signal

Jul 16, 2024

Looking for information on Verity SD100 monochromator unit

Jul 12, 2024

More from Optics

Related Stories

Graphite changes to hexagonal diamond in picoseconds

Aug 8, 2022

Microjets are faster than a speeding bullet

Dec 9, 2020

Study shows magnesium oxide undergoes dynamic transition when it comes to super-Earth exoplanets

Jun 12, 2024

Exploding electrical wires underwater to understand shock waves

May 2, 2019

Collapsing bubbles show new ejecta production mechanism can occur under multiple-shock conditions

Feb 22, 2023

Checking out iron under pressure

Jun 8, 2020

Recommended for you

X-ray imagery of vibrating diamond opens avenues for quantum sensing

Aug 7, 2024

Glimpse into the nanoworld: Microscope reveals tiniest cell processes

Hidden harmonies: Team discovers magnon–phonon Fermi resonance in an antiferromagnet

Coupling excitons to polaritons for better solar cells and higher intensity LEDs

Researchers develop energy-efficient optical neural networks

Aug 6, 2024

Researchers achieve super-Bloch oscillations in strong-driving regime

Aug 5, 2024

Let us know if there is a problem with our content

Use this form if you have come across a typo, inaccuracy or would like to send an edit request for the content on this page. For general inquiries, please use our contact form . For general feedback, use the public comments section below (please adhere to guidelines ).

Please select the most appropriate category to facilitate processing of your request

Thank you for taking time to provide your feedback to the editors.

Your feedback is important to us. However, we do not guarantee individual replies due to the high volume of messages.

E-mail the story

Your email address is used only to let the recipient know who sent the email. Neither your address nor the recipient's address will be used for any other purpose. The information you enter will appear in your e-mail message and is not retained by Phys.org in any form.

Newsletter sign up

Get weekly and/or daily updates delivered to your inbox. You can unsubscribe at any time and we'll never share your details to third parties.

More information Privacy policy

Donate and enjoy an ad-free experience

We keep our content available to everyone. Consider supporting Science X's mission by getting a premium account.

E-mail newsletter

Cracking and Precipitation Behavior of Refractory BCC–B2 Alloys Under Laser Melting Conditions

• Topical Collection: High Entropy Alloys
• Published: 06 August 2024

Cite this article

• Kaitlyn M. Mullin 1 ,
• Sebastian A. Kube 1 , 2 ,
• Sophia K. Wu 1 &
• Tresa M. Pollock 1  

16 Accesses

Explore all metrics

Emulating the Ni-base superalloy \(\gamma \)  +  \(\gamma ^{\prime }\) microstructure in BCC–B2 refractory alloys is a promising design strategy to achieve high temperature strength and ductility. Ru-base B2 precipitates have shown exceptional thermal stability but can be difficult to solutionize, making high cooling rate solidification pathways like additive manufacturing (AM) a promising approach for synthesis of more homogeneous microstructures. Using single track laser experiments on aged bulk substrates, five representative refractory alloys with varying Ru-base B2 precipitates (AlRu, HfRu, TiRu) and matrix constituents (Mo, Nb) were investigated for their solidification behavior and defect susceptibility under laser melting conditions. Susceptibility to solidification cracking, solid-state cracking, and keyhole formation was found to be highly dependent on the matrix composition. Characterization of the melt pools by scanning and transmission electron microscopy shows evidence for disordered BCC upon solidification, enabling tailoring of the B2 precipitates that are thermodynamically stable above 1300 °C. The B2 precipitate morphologies in the melt tracks after aging treatments are influenced by the partitioning behavior of Ru from laser melting. Results from these single track experiments provide guidance toward design strategies for fabricable refractory BCC–B2 alloys.

This is a preview of subscription content, log in via an institution to check access.

Access this article

Subscribe and save.

• Get 10 units per month
• Download Article/Chapter or eBook
• 1 Unit = 1 Article or 1 Chapter
• Cancel anytime

Price includes VAT (Russian Federation)

Instant access to the full article PDF.

Rent this article via DeepDyve

Institutional subscriptions

Similar content being viewed by others

Grain growth and Hall–Petch relationship in Ti37V15Nb22Hf23W3 refractory high-entropy alloys

Minor Elements and Solidification Cracking During Laser Powder-Bed Fusion of a High $$\gamma ^{\prime }$$ CoNi-Base Superalloy

Precipitate Rafting in a Polycrystalline Superalloy During Compression Creep

Data availability.

All data included in this study are available upon request by contact with the corresponding author.

T.M. Pollock and S. Tin: J. Propul. Power , 2012, vol. 22(2), pp. 361–74. https://doi.org/10.2514/1.18239 .

Article   Google Scholar  

T.M. Pollock: Nat. Mater. , 2016, vol. 15(8), 809–15. https://doi.org/10.1038/nmat4709 .

R.C. Reed: The Superalloys: Fundamentals and Applications , vol. 9780521859042, Cambridge University Press, Cambridge, 2006. https://doi.org/10.1017/CBO9780511541285 .

F. Schmidt, E. Bartlett, and H. Ogden: in AIME Proceedings , New York, 1962.

C. C. Wojcik: MRS Online Proc. Library (OPL) , 1993, vol. 322, pp. 519–30. https://doi.org/10.1557/PROC-322-519 .

R. Cooper, Jr. and E. Hoffman: Refractory alloy technology for space nuclear power applications, Tech. Rep., Oak Ridge, 1984.

J. Wadsworth: Miner. Met. Mater. Ser. , 2022. https://doi.org/10.1007/978-3-030-92381-5_23 .

D.B. Miracle and O.N. Senkov: Acta Mater. , 2017, vol. 122, pp. 448–511. https://doi.org/10.1016/J.ACTAMAT.2016.08.081 .

Article   CAS   Google Scholar  

O.N. Senkov, D.B. Miracle, K.J. Chaput, and J.P. Couzinie: J. Mater. Res. , 2018, vol. 33(19), pp. 3092–3128. https://doi.org/10.1557/JMR.2018.153 .

E.P. George, D. Raabe, and R.O. Ritchie: Rev. Mater. , 2019, vol. 4(8), pp. 515–34. https://doi.org/10.1038/s41578-019-0121-4 .

S.A. Kube, S. Sohn, D. Uhl, A. Datye, A. Mehta, and J. Schroers: Acta Mater. , 2019, vol. 166, pp. 677–86. https://doi.org/10.1016/J.ACTAMAT.2019.01.023 .

D. Miracle, B. Majumdar, K. Wertz, and S. Gorsse: Scripta Mater. , 2017, vol. 127, pp. 195–200. https://doi.org/10.1016/J.SCRIPTAMAT.2016.08.001 .

D.B. Miracle, M. H. Tsai, O. N. Senkov, V. Soni, and R. Banerjee: Scripta Mater. , 2020, vol. 187, pp. 445–52. https://doi.org/10.1016/J.SCRIPTAMAT.2020.06.048 .

E.A. Lass: Metall. Mater. Trans. A , 2022. https://doi.org/10.1007/S11661-022-06844-6 .

O.N. Senkov, J.K. Jensen, A.L. Pilchak, D.B. Miracle, and H.L. Fraser: Mater. Des. , 2018, vol. 139, pp. 498–511. https://doi.org/10.1016/J.MATDES.2017.11.033 .

J.P. Couzinié, M. Heczko, V. Mazánová, O.N. Senkov, M. Ghazisaeidi, R. Banerjee, and M.J. Mills: Acta Mater. , 2022, vol. 233, p. 117995. https://doi.org/10.1016/J.ACTAMAT.2022.117995 .

O.N. Senkov, B. Crossman, S.I. Rao, J.P. Couzinie, D.B. Miracle, T.M. Butler, R. Banerjee, and M. Mills: Acta Mater. , 2023, vol. 254, p. 119017. https://doi.org/10.1016/J.ACTAMAT.2023.119017 .

T.E. Whitfield, G.J. Wise, H.J. Stone, and N.G. Jones: Appl. Phys. Lett. , 2021, vol. 119(21), p. 211901. https://doi.org/10.1063/5.0068045 .

T.E. Whitfield, N.L. Church, H.J. Stone, and N.G. Jones: J. Alloys Compd. , 2023, vol. 939, p. 168369. https://doi.org/10.1016/J.JALLCOM.2022.168369 .

N. Parkes, R. Dodds, A. Watson, D. Dye, C. Hardie, S.A. Humphry-Baker, and A.J. Knowles: Int. J. Refract. Met. Hard Mater. , 2023, vol. 113, p. 106209. https://doi.org/10.1016/J.IJRMHM.2023.106209 .

K. Ma, T. Blackburn, J.P. Magnussen, M. Kerbstadt, P.A. Ferreirós, T. Pinomaa, C. Hofer, D.G. Hopkinson, S.J. Day, P.A. Bagot, M.P. Moody, M.C. Galetz, and A.J. Knowles: Acta Mater. , 2023, vol. 257, p. 119183. https://doi.org/10.1016/J.ACTAMAT.2023.119183 .

E. Panina, N. Yurchenko, G. Salishchev, S. Zherebtsov, and N. Stepanov: Int. J. Refract. Met. Hard Mater. , 2023, vol. 115, p.106297. https://doi.org/10.1016/J.IJRMHM.2023.106297 .

N. Yurchenko, E. Panina, Rogal, L. Shekhawat, S. Zherebtsov, and N. Stepanov: Mater. Res. Lett. , 2023, vol. 10(2), pp. 78–87. https://doi.org/10.1080/21663831.2021.2022033 .

C. Frey, R. Silverstein, and T.M. Pollock: Acta Mater. , 2022, vol. 229, p. 117767. https://doi.org/10.1016/J.ACTAMAT.2022.117767 .

C. Frey, H. You, S. Kube, G.H. Balbus, K. Mullin, S. Oppenheimer, C.S. Holgate, and T.M. Pollock: Metall. Mater. Trans. A , 2024, vol. 55A, pp. 1739–64. https://doi.org/10.1007/s11661-024-07368-x .

S.A. Kube, C. Frey, C. McMullin, B. Neuman, K.M. Mullin, and T.M. Pollock: Acta Mater. , 2024, vol. 26, p. 119628. https://doi.org/10.1016/J.ACTAMAT.2023.119628 .

N.R. Philips, M. Carl, and N.J. Cunningham: Metall. Mater. Trans. A , 2020, vol. 51, pp. 3299–3310. https://doi.org/10.1007/s11661-020-05803-3 .

A. Talignani, R. Seede, A. Whitt, S. Zheng, J. Ye, I. Karaman, M.M. Kirka, Y. Katoh, and Y.M. Wang: Addit. Manuf. , 2002, vol. 58, p. 103009. https://doi.org/10.1016/J.ADDMA.2022.103009 .

P. Fernandez-Zelaia, C. Ledford, E.A. Ellis, Q. Campbell, A.M. Rossy, D.N. Leonard, and M.M. Kirka: Mater. Des. , 2021, vol. 207, p. 109809. https://doi.org/10.1016/J.MATDES.2021.109809 .

P. Rebesan, M. Ballan, M. Bonesso, A. Campagnolo, S. Corradetti, R. Dima, C. Gennari, G.A. Longo, S. Mancin, M. Manzolaro, G. Meneghetti, A. Pepato, E. Visconti, and M. Vedani: Addit. Manuf. , 2021, vol. 47, p. 102277. https://doi.org/10.1016/J.ADDMA.2021.102277 .

M. Higashi and T. Ozaki: Mater. Des. , 2020, vol. 191, p. 108588. https://doi.org/10.1016/J.MATDES.2020.108588 .

L. Thijs, M.L. Montero Sistiaga, R. Wauthle, Q. Xie, J.P. Kruth, and J. Van Humbeeck: Acta Mater. , 2012, vol. 61(12), pp. 4657–68. https://doi.org/10.1016/J.ACTAMAT.2013.04.036 .

V. Livescu, C.M. Knapp, G. T. Gray, R.M. Martinez, B.M. Morrow, and B.G. Ndefru: Materialia , 2018, vol. 1, pp. 15–24. https://doi.org/10.1016/J.MTLA.2018.06.007 .

P.D. Awasthi, P. Agrawal, R.S. Haridas, R.S. Mishra, M.T. Stawovy, S. Ohm, and A. Imandoust: Mater. Sci. Eng. A , 2022, vol. 831, p. 142183. https://doi.org/10.1016/J.MSEA.2021.142183 .

S.S. Babu, N. Raghavan, J. Raplee, S.J. Foster, C. Frederick, M. Haines, R. Dinwiddie, M.K. Kirka, A. Plotkowski, Y. Lee, and R.R. Dehoff: Metall. Mater. Trans. A , 2018, vol. 49A, pp. 3764–80. https://doi.org/10.1007/S11661-018-4702-4 .

T. DebRoy, H.L. Wei, J. S. Zuback, T. Mukherjee, J.W. Elmer, J.O. Milewski, A.M. Beese, A. Wilson-Heid, A. De, and W. Zhang: Prog. Mater. Sci. , 2018, vol. 92, pp. 112–224. https://doi.org/10.1016/J.PMATSCI.2017.10.001 .

B. Graybill, M. Li, D. Malawey, C. Ma, J.M. Alvarado-Orozco, and E. Martinez-Franco: in ASME 2018 13th International Manufacturing Science and Engineering Conference, MSEC 2018 , 2018, vol. 1(9). https://doi.org/10.1115/MSEC2018-6666 .

M. Prager and C. Shira: Weld. Res. Council Bull. , 1968, vol. 6, p. 128.

D.G. Eskin and L. Katgerman: Metall. Mater. Trans. A , 2007, vol. 38A(7), pp. 1511–19. https://doi.org/10.1007/S11661-007-9169-7 .

S. Kou: JOM , 2003, vol. 55(6), pp. 37–42. https://doi.org/10.1007/s11837-003-0137-4 .

M.C. Brennan, J.S. Keist, T.A. Palmer, D.L. Bourell, W. Frazier, H. Kuhn, and M. Seifi: J. Mater. Eng. Perform. , 2021, vol. 30(7), pp. 4808–18. https://doi.org/10.1007/S11665-021-05919-6 .

I. Lopez-Galilea, B. Ruttert, J. He, T. Hammerschmidt, R. Drautz, B. Gault, and W. Theisen: Addit. Manuf. , 2019, vol. 30, p. 100874. https://doi.org/10.1016/J.ADDMA.2019.100874 .

T.M. Pollock, A.J. Clarke, and S.S. Babu: Metall. Mater. Trans. A , 2020, vol. 51A, pp. 6000–19. https://doi.org/10.1007/S11661-020-06009-3/FIGURES/13 .

S.P. Murray, K.M. Pusch, A.T. Polonsky, C.J. Torbet, G.G. Seward, N. Zhou, S.A. Forsik, P. Nandwana, M.M. Kirka, R.R. Dehoff, W.E. Slye, and T.M. Pollock: Nat. Commun. , 2020, vol. 11(1), pp. 1–11. https://doi.org/10.1038/s41467-020-18775-0 .

Y.T. Tang, C. Panwisawas, J.N. Ghoussoub, Y. Gong, J.W. Clark, A.A. Németh, D.G. McCartney, and R.C. Reed: Acta Mater. , 2021, vol. 202, pp. 417–36. https://doi.org/10.1016/J.ACTAMAT.2020.09.023 .

B. Vrancken, R.K. Ganeriwala, and M.J. Matthews: Acta Mater. , 2020, vol. 194, pp. 464–72. https://doi.org/10.1016/J.ACTAMAT.2020.04.060 .

G. Lessmann: Determination of Weldability and Elevated Temperature Stability of Refractory Metal Alloys: I—Weldability of Refractory Metal Alloys. Tech. Rep., NASA Contractor Report, Pittsburgh, PA, 1970.

M.H. Scott and P.M. Knowlson: J. Less Common Met. , 1963, vol. 5(3), pp. 205–44. https://doi.org/10.1016/0022-5088(63)90029-8 .

B.A. Chin, P. Banerjee, J. Liu, and S. Chen: Weld. Brazing Solder. , 1993. https://doi.org/10.31399/ASM.HB.V06.A0001426 .

K.M. Mullin, C. Frey, J. Lamb, S.K. Wu, M.P. Echlin, and T.M. Pollock: Mater. Des., 2024, vol. 238, p. 112726. https://doi.org/10.1016/J.MATDES.2024.112726 .

J. Stephens and W. Witzke: (NASA-TN-D-7847) Hardness behavior of binary and ternary niobium alloys at 77 and 300 K. Tech. rep., Lewis Research Center, NASA, 1974.

P. Suárez Ocaño, A. Manzoni, I. Lopez-Galilea, B. Ruttert, G. Laplanche, and L. Agudo Jácome: J. Alloys Compd. , 2023, vol. 949(2023), p. 169871. https://doi.org/10.1016/J.JALLCOM.2023.169871 .

A. Chesetti, S. Banerjee, S. Dasari, M.S.K. Nartu, S.M. Varahabhatla, A. Sharma, A. Ramakrishnan, D. Satko, S. Gorsse, A. Salem, and R. Banerjee: Scripta Mater. , 2023, vol. 225, p. 115160. https://doi.org/10.1016/J.SCRIPTAMAT.2022.115160 .

G.G. Gladush and I. Smurov: Physics of Laser Materials Processing: Theory and Processing , vol. 146, Springer Series in Materials Science, Springer, 2011.

Z. Gan, O.L. Kafka, N. Parab, C. Zhao, L. Fang, O. Heinonen, T. Sun, and W.K. Liu: Nat. Commun. , 2021, vol. 12(1), pp. 1–8. https://doi.org/10.1038/s41467-021-22704-0 .

Z.Y. Wang, J.T. Liu, D.M. Hirak, D.C. Weckman, and H.W. Kerr: J. Laser Appl. , 1993, vol. 5(1), pp. 5–12. https://doi.org/10.2351/1.4745318 .

D.B. Williams and C.B. Carter: Transmission Electron Microscopy: A Textbook for Materials Science , 2nd ed., Springer New York, 2009.

Book   Google Scholar  

K.M. Bertsch, G. Meric de Bellefon, B. Kuehl, and D.J. Thoma: Acta Mater. , 2020, vol. 199, pp. 19–33. https://doi.org/10.1016/J.ACTAMAT.2020.07.063 .

S. Tekumalla, M. Seita, and S. Zaefferer: Acta Mater. , 2024, vol. 262, p. 119413. https://doi.org/10.1016/J.ACTAMAT.2023.119413 .

D. Kong, C. Dong, S. Wei, X. Ni, L. Zhang, R. Li, L. Wang, C. Man, and X. Li: Addit. Manuf. , 2021, vol. 38, p. 101804. https://doi.org/10.1016/J.ADDMA.2020.101804 .

F. Cao, T.K. Nandy, D. Stobbe, and T.M. Pollock: Intermetallics , 2007, vol. 15(1), pp. 34–43. https://doi.org/10.1016/J.INTERMET.2006.03.004 .

P. Bajaj, J. Wright, I. Todd, and E.A. Jägle: Addit. Manuf. , 2019, vol. 27, pp. 246–58. https://doi.org/10.1016/J.ADDMA.2018.12.003 .

A.K. Agrawal, B. Rankouhi, and D.J. Thoma: Curr. Opin. Solid State Mater. Sci. , 2022, vol. 26(6), p. 101024. https://doi.org/10.1016/J.COSSMS.2022.101024 .

C.L.A. Leung, S. Marussi, R.C. Atwood, M. Towrie, P.J. Withers, and P.D. Lee: Nat. Commun. , 2018, vol. 9(1), pp. 1–9. https://doi.org/10.1038/s41467-018-03734-7 .

C.J. Smith, S. Tammas-Williams, E. Hernandez-Nava, and I. Todd: Sci. Rep. , 2017, vol. 7(1), pp. 1–8. https://doi.org/10.1038/s41598-017-11243-8 .

S.A. Khairallah, A.M. Rubenchik, and C.D. Boley: Appl. Opt. , 2015, vol. 54(9), pp. 2477–82. https://doi.org/10.1364/AO.54.002477 .

Article   CAS   PubMed   Google Scholar  

H. Esaka and W. Kurz: J. Cryst. Growth , 1985, vol. 72(3), pp. 578–84. https://doi.org/10.1016/0022-0248(85)90208-8 .

J.L. Bartlett and X. Li: Addit. Manuf. , 2019, vol. 27, pp. 131–49. https://doi.org/10.1016/J.ADDMA.2019.02.020 .

A. Miklas: Additive Manufacturing of Refractory Alloys, Masters thesis, Colorado School of Mines, 2022.

J.L. Murray: Bull. Alloy Phase Diagrams , 1982, vol. 3(2), pp. 216–21. https://doi.org/10.1007/BF02892392 .

P. Promoppatum, S.C. Yao, P.C. Pistorius, and A.D. Rollett: Engineering , 2017, vol. 3(5), pp. 685–94. https://doi.org/10.1016/J.ENG.2017.05.023 .

J. Yoshioka and M. Eshraghi: Heat Mass Transf. , 2023, vol. 59(7), pp. 1155–66. https://doi.org/10.1007/S00231-022-03318-8 .

S. Kou: Welding Metallurgy , 3rd ed., Wiley, New York, 2020.

T.W. Clyne, M. Wolf, and W. Kurz: Metall. Trans. B , 1982, vol. 13(2), pp. 259–66. https://doi.org/10.1007/BF02664583 .

E.B. Raeker, K.M. Pusch, S.A. Forsik, N. Zhou, A.D. Dicus, Q.Q. Ren, J.D. Poplawsky, M.M. Kirka, and T.M. Pollock: Metall. Mater. Trans. A , 2023, vol. 54A, pp. 1744–57. https://doi.org/10.1007/S11661-023-06957-6 .

J.N. Ghoussoub, Y.T. Tang, W.J. Dick-Cleland, A.A. Németh, Y. Gong, D.G. McCartney, A.C. Cocks, and R.C. Reed: Metall. Mater. Trans. A , 2022, vol. 53A, pp. 962–83. https://doi.org/10.1007/S11661-021-06568-Z .

R. Puerling, A. Miklas, F.G. Coury, N.R. Philips, P. Mason, N.E. Peterson, A. Deal, J. Klemm-Toole, and A.J. Clarke: Metall. Mater. Trans. A , 2023, vol. 54(4) , pp. 1070–76. https://doi.org/10.1007/s11661-023-06980-7 .

O.N. Senkov, G.B. Wilks, D.B. Miracle, C.P. Chuang, and P.K. Liaw: Intermetallics , 2010, vol. 18(9), pp. 1758–65. https://doi.org/10.1016/J.INTERMET.2010.05.014 .

L.H. Mills, M.G. Emigh, C.H. Frey, N.R. Philips, S.P. Murray, J. Shin, D.S. Gianola, and T.M. Pollock: Acta Mater. , 2023, vol. 245, p. 118618. https://doi.org/10.1016/J.ACTAMAT.2022.118618 .

C.K. Borg, C. Frey, J. Moh, T.M. Pollock, S. Gorsse, D.B. Miracle, O.N. Senkov, B. Meredig, and J.E. Saal: Sci. Data , 2020, vol. 7(1), pp. 1–6. https://doi.org/10.1038/s41597-020-00768-9 .

A. Biesiekierski, J. Lin, Y. Li, D. Ping, Y. Yamabe-Mitarai, and C. Wen: Acta Biomater. , 2017, vol. 48, pp. 461–67. https://doi.org/10.1016/J.ACTBIO.2016.09.012 .

J.L. Murray: Bull. Alloy Phase Diagrams , 1981, vol. 2(2), pp. 185–92. https://doi.org/10.1007/BF02881476 .

Y. Zhang, H. Liu, and Z. Jin: Calphad , 2001, vol. 25, pp. 305–317. https://doi.org/10.1016/S0364-5916(01)00051-7 .

J. Liu, C. Zhou, and H. Wang: Calphad , 2021, vol. 72, p. 102239. https://doi.org/10.1016/J.CALPHAD.2020.102239 .

H. Okamoto: in Alloy Phase Diagrams , ASM International, Materials Park, 2016, pp. 89–89. https://doi.org/10.31399/ASM.HB.V03.A0006247 .

M.M. Kirka, Y. Lee, D.A. Greeley, A. Okello, M.J. Goin, M.T. Pearce, and R.R. Dehoff: JOM , 2017, vol. 69(3), pp. 523–31. https://doi.org/10.1007/S11837-017-2264-3 .

Download references

Acknowledgments

This research was developed with funding from the Office of Naval Research (ONR) under Grants N00014-181-2392 and N00014-22-1-2087. KM was supported by a National Science Foundation (NSF) Graduate Research Fellowship under Grant No. 2139319. The research reported here made use of the shared facilities of the Materials Research Science and Engineering Center (MRSEC) at UC Santa Barbara: NSF DMR-2308708. The UC Santa Barbara MRSEC is a member of the Materials Research Facilities Network ( www.mrfn.org ). The authors gratefully acknowledge Ben Neuman and Chiyo McMullin for their assistance with material preparation and compression testing, and Carolina Frey for discussions on refractory BCC–B2 alloys.

Author information

Authors and affiliations.

Materials Department, University of California Santa Barbara, UCSB Engineering II 1355, Santa Barbara, 93106, CA, USA

Kaitlyn M. Mullin, Sebastian A. Kube, Sophia K. Wu & Tresa M. Pollock

Department of Materials Science and Engineering, University of Wisconsin - Madison, Madison, 53706, WI, USA

Sebastian A. Kube

You can also search for this author in PubMed   Google Scholar

Corresponding author

Correspondence to Kaitlyn M. Mullin .

Ethics declarations

Conflict of interest.

The authors declare that they have no conflict of interest.

Additional information

Publisher's note.

Springer Nature remains neutral with regard to jurisdictional claims in published maps and institutional affiliations.

Rights and permissions

Springer Nature or its licensor (e.g. a society or other partner) holds exclusive rights to this article under a publishing agreement with the author(s) or other rightsholder(s); author self-archiving of the accepted manuscript version of this article is solely governed by the terms of such publishing agreement and applicable law.

Reprints and permissions

About this article

Mullin, K.M., Kube, S.A., Wu, S.K. et al. Cracking and Precipitation Behavior of Refractory BCC–B2 Alloys Under Laser Melting Conditions. Metall Mater Trans A (2024). https://doi.org/10.1007/s11661-024-07541-2

Download citation

Received : 29 March 2024

Accepted : 25 July 2024

Published : 06 August 2024

DOI : https://doi.org/10.1007/s11661-024-07541-2

Share this article

Anyone you share the following link with will be able to read this content:

Sorry, a shareable link is not currently available for this article.

Provided by the Springer Nature SharedIt content-sharing initiative

Advertisement

• Find a journal
• Publish with us
• Track your research