Since the construction of the first wind turbine in 1991, many technological advancements have made the utilization of offshore wind energy a reality. According to 2023’s Global Offshore Wind Report, 2022 was the second-highest year in history for offshore installations, which brought 8.8 GW of new offshore wind into the grid, bringing the total offshore wind power capacity to 64.3 GW. That represents 7.1 percent of global wind installation. It has been projected that by 2050, offshore wind will contribute about 34 percent of total wind production.
With the recent innovations into floating offshore wind to support the growth of green energy, turbines can now be installed in deeper and more complex seabed locations. This has been made possible by floating platforms anchored to heavy weights on the seabed through flexible steel cables and chains. The four different types of offshore floating platforms that have currently been developed allow for adaptability of this technology depending on the specific conditions of each site.
With the recent innovations into floating offshore wind to support the growth of green energy, turbines can now be installed in deeper and more complex seabed locations. (Courtesy: NatureMetrics)
These floating platforms are designed to give turbines buoyancy in diverging seabed conditions and wind types and can extend manufacturing capability while reducing the cost of floating wind farms. As a result, floating offshore wind is expected to achieve a total capacity of 380 GW by 2030, according to the Global Offshore Wind Alliance (GOWA).
Global interest in floating offshore wind has been growing exponentially, as countries aim to reach renewable energy targets by 2030 and beyond, to reach net zero. In March of last year, White House National Climate Adviser Ali Zaidi announced the release of the U.S.’ very own Offshore Wind Energy Strategy. Through this, the country is aiming to leverage all potential resources to harness offshore wind energy as a reliable energy source and create thousands of well-paying union jobs that will simultaneously help to revitalize coastal communities and position the U.S. as a leader in offshore wind energy. With this as a backdrop, the Biden administration set a goal to generate 30 GW of floating offshore wind energy by 2030, which would power more than 10 million homes.
Although investment into engineering technologies allowed for European renewable energy to be price-competitive with conventional power sources in 2017, the U.S. has gone a step further and set the Floating Offshore Wind Shot goal of reducing the cost of offshore wind by more than 70 percent, which would further aid in the positioning as the global sector leader. In becoming a global leader, the U.S. geography lends itself to the use of various offshore technologies. Where the East Coast is opting for traditional piled-soiled platforms, the deeper waters of the West Coast make floating wind the only viable option if offshore wind is to be introduced.
Innovative approaches such as environmental DNA (eDNA) analysis and earth observation are enabling offshore wind companies to accurately measure and track the environmental impact of business operations to boost productivity that are inherently tied to the environment in which they operate. (Courtesy: NatureMetrics)
Environmental impact and regulation
Despite the floating offshore wind technology being nascent with most projects currently piloted or at the pre-commercial stage, the total pipeline of projects is growing, and floating wind is expected to reach 15 percent of total energy produced by offshore wind by 2050. However, little is yet known about the impact this emerging technology might have on marine ecosystems. The technical developments of floating offshore winds will allow for the placement of these floating turbines in deep-sea areas that have very rich habitats, so it will be critical to understand how the technology will affect these important environments.
Targets to guide each country’s marine regulatory efforts have already been suggested within the Kumming-Montreal Global Biodiversity Framework, which set the goal to recover and restore 30 percent of global marine ecosystems, halting human-induced extinction of threatened species through the implementation of conservation measures and the sustainable use of seascapes and the ocean. To adhere to these targets, there must be a just and sustainable deployment of floating offshore wind to go hand-in-hand with the advancements in design and manufacturing that will allow for this renewable energy to achieve its full potential.
Traditional methods such as trawling have historically been used to track the impact of offshore wind farms. However, these methods require specialized vessels and ecologists, are dependent on the weather conditions at sea, and netting techniques are unable to be used within wind-turbine arrays leading to a lack of biodiversity data at the closest affected site. New technologies developed in recent years have allowed for the development of sampling methods that provide a more accurate picture of the biodiversity in the environment while tackling many of the issues that traditional methods present.
Nature Intelligence Technologies: An Essential Tool
Innovative approaches such as environmental DNA (eDNA) analysis and earth observation are enabling offshore wind companies to accurately measure and track the environmental impact of business operations to boost productivity that are inherently tied to the environment in which they operate; eDNA technology, which can detect traces of DNA that species leave behind in their environment, allows for the identification of up to 70 percent more species from a water or sediment sample. Earth observation methods such as satellites and drone remote sensing technologies allow for the visual monitoring of habitat types, distribution of species, and even the creation of digital twins replicating an entire ecosystem to assess biodiversity. The use of these nature intelligence technologies will allow offshore wind companies to build a complete map of the ecosystems in which they operate and future-proof marine construction projects against regulatory frameworks.
Some companies are already successfully using these technologies, which have proven to uncover a greater diversity of species than previously thought given the information that had been collected through traditional methods. With funding from the U.K.’s Offshore Wind Growth Partnership’s Innovation Grant, EDF Renewables’ Blyth Offshore Demonstrator, a local coastal wind farm half a mile off the coast of Blyth, Northumberland, England, set out to investigate whether eDNA sampling was an effective and replicable methodology to the standardized trawling method — which is invasive to the analyzed ecosystem. The sampling involved the collection of water samples one meter above the seabed, reducing the safety risks that trawling entails, and personnel costs as this sampling can be conducted as part of regular site works.
The technical developments of floating offshore winds will allow for the placement of these floating turbines in deep-sea areas that have very rich habitats, so it will be critical to understand how the technology will affect these important environments. (Courtesy: NatureMetrics)
The eDNA method was not only able to detect more species compared to the traditional trawl method, but it was also able to detect mammals, invertebrates, and birds. In addition, because the eDNA testing is done with a simple water sample, this enabled surveying to be done within the turbine array, where trawling cannot be conducted. This led to data insights within the array that showed greater relative abundance of reef-dwelling and commercially important fish species, which supports a hypothesis that the structures provide an artificial reef and nursery environment. The specificity of insights these technologies bring to the table will be increasingly valuable and standardized due to the rise in frameworks informing marine conservation.
As communities continue to need more energy and green energy becomes a preferred choice, we will only see more innovation within the energy transition era. But making sure we’re protecting ecosystems we’re trying to save by using green energy is imperative. Companies will need to be equally innovative in how they measure nature to navigate inefficiencies and costs, while also capturing robust data on the ecosystems they affect, as they continue to deploy early-stage technologies in the field; eDNA is helping to provide an answer to these challenges, ultimately supporting the green transition while also ensuring a sustainable foundation for energy businesses and a brighter future for our planet.
The construction of the SeAH Wind Monopile Factory will create 1,500 jobs during the supply chain and construction phases, and an additional 750 jobs once fully operational in 2026. This project, in the Teesworks industrial zone, is the largest monopile factory in the world, spanning a 90-acre site. Its main structure is 800 meters long and has the capacity to produce monopiles weighing up to 3,000 tons, essential components for installing offshore wind turbines.
Sarens, global leader and reference in crane rental services, heavy lifting and engineered transport, is involved in the construction of the factory on behalf of its customer Severfield UK and the owner of the facility, SeAH.
To perform the diverse tasks at the facility, the Sarens engineering team selected a range of cranes, including seven LTR1100 units, with a load capacity of up to 100 tons, seven LR130 units, one LTM1750, two LTM1650s, which can lift up to 650 tons in a single load, as well as a CC2800 crawler crane. Sarens is using these cranes to lift structural parts and steel frames, which are being moved on site by self propelled modular transport.
The SeAH Wind monopile factory will play a crucial role in reducing the global carbon footprint through the production of components for offshore wind energy, one of the cleanest and most sustainable forms of power generation.
SeAH Wind has already established strategic agreements with offshore wind-power companies, such as Ørsted, to become the first and main customer for the monopile foundations manufactured at the U.K. facility. These partnerships are essential to ensure the long-term viability of the factory and to strengthen the offshore wind supply chain.
One of the main challenges of the job consists of the lifting site being exposed and susceptible to high winds. To address this issue, Sarens worked with the crane manufacturer to obtain a wind-speed increase based on the crane configuration, lifting criteria (radius, capacity, etc.), and load characteristics (sail area, weight, load coefficient).
Sarens has extensive international experience in the assembly and maintenance of wind farms. It has participated in installations around the world and particularly in Europe, as in France (Saint Nazaire and Saint Brieuc) and the U.K., where its last project is now successfully completed.
Recently, Sarens has worked in the marshalling of 62 of the monopiles, each weighing 2,000 metric tons, the largest and heaviest XXL monopiles ever to be handled in the U.K., and now the 882-MW Moray West offshore wind farm, is well on its way to contribute to the Scottish renewable energy network.
Vestas recently secured an 81-MW order from Invenergy for the wind-energy project Pencloe in Dumfries and Galloway in Scotland. Vestas will deliver 18 V136-4.5 MW wind turbines, and the order includes supply, delivery, installation, and commissioning of the turbines. Upon completion, Vestas will service the turbines under a multi-year Active Output Management 5000 (AOM 5000) service agreement designed to ensure assets’ performance.
Vestas will deliver 18 wind turbines to the sites in the U.K. (Courtesy: Vestas)
“We are delighted to achieve this key milestone in our collaboration with Invenergy,” said James Ian Robinson, country manager and director Sales UK for Vestas Northern and Central Europe.
“Vestas’ technology delivers a robust business case for the competitive U.K. electricity market. We look forward to the execution phase where we will continue our strong construction track record. We thank Invenergy for trusting Vestas with their largest wind park to date in the U.K.”
“We are excited to be utilizing state-of-the-art Vestas turbines at the Pencloe Wind Energy Centre, which will be the largest Invenergy-developed project in the United Kingdom,” said Stuart Winter, vice president and country manager at Invenergy. “This project not only represents our mission to accelerate cleaner, more reliable and affordable energy, but also underscores our dedication to fostering positive community relationships and ensuring local economies benefit from our projects.”
Turbine delivery is expected to begin in the second quarter of 2025 with the project expected to be fully operational in early 2026.
Revolution Wind is marking the completed construction of the project’s union-built advanced foundation components, the latest milestone for Rhode Island and Connecticut’s first large-scale offshore wind farm and the first multi-state offshore wind farm in the nation.
Rhode Island also welcomed the arrival of the offshore wind service operations vessel, the Eco Edison. “Revolution Wind is a win for Rhode Island’s environment and our economy,” said Gov.Dan McKee.
The Eco Edison vessel will serve as a floating, year-round home base for turbine technicians. (Courtesy: Revolution Wind)
“We’re excited by the progress of this project, which is supporting good-paying jobs and propelling our state toward a stronger blue economy and a more sustainable future.” Ørsted and Eversource’s Revolution Wind is directly creating roughly 1,200 jobs across Rhode Island and Connecticut and accelerating the states’ clean-energy sectors with investments in workforce development, union partnerships, shipbuilding, and port infrastructure.
“I am delighted to participate in (the) event, which highlights the tremendous progress that the Biden-Harris Administration has made, along with our state and industry partners, toward achieving our goal of deploying 30 GW of offshore wind energy capacity by 2030,” said BOEM Director Elizabeth Klein. “Today represents another step forward in providing clean-energy jobs and investing in our coastal communities while strengthening America’s energy security.”
Revolution Wind will have the capacity to generate 400 MW of affordable offshore wind power for Rhode Island and 304 MW of the same for Connecticut, enough clean energy to power more than 350,000 homes across both states and bring each closer to reaching their climate targets. The Eco Edison, which will be based out of ProvPort during Revolution Wind’s construction is an example of Ørsted’s $20 billion investment in building out an American clean-energy industry.
This first-ever American-built, owned, and crewed offshore wind service operations vessel will serve as a floating, year-round home base for the turbine technicians. These techs – Rhode Islanders among them – will work at sea over the life of the wind farms, servicing and maintaining the wind turbines. The vessel will play an integral part of the operation and maintenance of Ørsted and Eversource’s Northeast projects, using Port of Quonset, Rhode Island, as well as Port Jefferson, New York.
“Rhode Island is the birthplace of American offshore wind, and the state is continuing to harness the true potential of offshore wind to transform its ports, workforce, and economy,” said David Hardy, Group EVP and CEO Americas at Ørsted. “Thanks to our local union and supply-chain partners and our talented construction team, we’re building and delivering Revolution Wind.And it’s only fitting that the Ocean State will host our state-of-the-art, American-made service vessel, the Eco Edison, during the construction of this historic project for New England.”
The Revolution Wind project site, about 15 miles south of the Rhode Island coast and 32 miles southeast of the Connecticut coast, is adjacent to Ørsted and Eversource’s South Fork Wind, America’s first utility-scale offshore wind farm. The site is expected to be in operation in 2025.
TEPCO Renewable Power and Shoreline Wind recently announced a collaboration aimed at optimizing the design and management of offshore wind-power operations and maintenance (O&M) by using intelligent simulation software.
TEPCO Renewable Power, a leader in Japan’s renewable energy sector, is on a mission to expand its offshore wind portfolio. TEPCO determined Shoreline Wind’s O&M simulation software is an ideal tool to solution to evaluate maintenance and management costs, predict future cash flows, and streamline operational logistics.
TEPCO and Shoreline Wind are collaborating on using simulation software. (Courtesy: Shoreline Wind)
By integrating Shoreline Wind’s O&M software into the design phase of its offshore wind projects, TEPCO is supported with a digital framework for insights into costs, availability, and resource use during the operations and maintenance campaigns. A key task for TEPCO is the estimation of maintenance costs and vessel specifications, including the costs and specifications of Crew Transfer Vessels during operations.
TEPCO’s collaboration with Shoreline Wind marks a significant step toward more innovative and digitalized offshore wind operations. By using data-driven outputs strategically for O&M planning, TEPCO Renewable Power is supporting its offshore wind operations with reduced uncertainties and improved reliability.
What is your role with Locke Lord and your interest in the offshore wind energy sector?
I’m senior counsel at Locke Lord, a premier full-service law firm. I advise offshore wind developers, advocacy groups, and companies in the supply chain on legal and policy issues. I also represent onshore renewable energy companies on permitting issues.
I’ve been involved in offshore wind law and policy for well over a decade. For five years, I was a career attorney at the solicitor’s office at the Department of Interior where I was frontline attorney adviser to the Bureau of Ocean Energy Management’s (BOEM) offshore wind program straddling the Obama and Trump administrations.
I spent a year and a half at GE as commercial counsel for their U.S. offshore wind business, and then two years at the American Clean Power Association as their vice president for offshore wind before coming to Locke Lord.
BOEM recently released the next phase of its Renewable Energy Modernization Rule. Could you explain what this means for the future of U.S. offshore wind?
The U.S. offshore wind industry is facing some positive trends and some challenges for sure. The modernization rule plays an unequivocally positive role, although there were some missed opportunities for it to be even more of a game changer.
First off, I agree with BOEM that the rule will save developers a sizable amount of money. Obviously, when you consider the massive capital costs for these projects, these savings may not seem huge — but they are not insignificant.
But perhaps just as important as the cost savings, the rule creates efficiencies and codifies regulatory practices that had previously been done on more of an ad-hoc basis. Developers, particularly in a complex industry like this, need certainty. Having that certainty codified in the regulations is crucial.
In turn, that makes the industry more resilient in the face of economic and political uncertainty — that everyone understands what the rules are and that they align with how the industry actually functions.
What brought about the need for this updated rule?
In a nutshell, the rule is about the collective need for alignment between regulations and industry practice. The Department of Interior got the authority to regulate offshore wind in 2005, and then in 2009, it promulgated offshore wind regulations that remained virtually unchanged for 15 years. When you think back to 2009, there certainly weren’t any offshore wind farms spinning off of the U.S., and there was only one proposal in federal waters, Cape Wind. The writers of the regulations didn’t know nearly as much about how offshore wind works as BOEM does today.
So instead, they based their regulations off of what they knew, which was offshore oil and gas primarily in the Gulf of Mexico. Given the lack of industry knowledge at the time, the fact that U.S. offshore wind has gotten as far as it has with those regulations is actually quite impressive. One of the main reasons is that the original drafters had the foresight to include a mechanism to allow for case-by-case exemptions to the rules if they weren’t functioning well.
Over time, it became clear there were aspects of the rules that just didn’t align with industry practices or simply weren’t efficient. These issues were aggregated and evolved into the Modernization Rule.
Getting into some of the specifics of the rule, how will it help projects better align with the European practices, and why is this important?
It’s vital for the U.S. and other countries to look to Europe for lessons because they have the most mature industry and strong environmental laws. BOEM and BSEE, to their credit, have done that with the Mod Rule.
The biggest example of this is BOEM’s decision to codify the project design envelope (PDE), which is the practice of proposing a range of design parameters for the project in your construction operation plan (COP). As the project moves through the permitting process, the developer gets more commercial information and works with key stakeholders and can narrow down the project parameters until it reaches a final project proposal. And so BOEM is saying in the Mod Rule: “Yes, we want you to use the project design envelope approach.” This ensures that, as developers refine their PDE, it doesn’t trigger a completely new environmental analysis, require resubmittal of the COP, and unduly delay the permitting process.
Another key change that looks to the European experience is the timing of providing financial assurance for decommissioning. Decommissioning costs are significant, and it makes sense that developers would provide security to ensure money is available to decommission projects when they reach the end of their lifespan. The original regulations had been interpreted to require that the entire decommissioning financial assurance be provided up front.
In Europe, it’s provided on more of a schedule, and so that’s what BOEM has adopted. This will save developers a lot of money by not locking up as much capital on the front end, while still ensuring the decommissioning costs will be covered when they’re needed.
What kind of options will the rule give developers looking to increase onshore fabrication of project components?
The rule as it was originally drafted suggested that no fabrication could take place until you were at the very end of the permitting process and all your engineering reports had been reviewed and there were no objections. This is a big problem. Developers must be able to start their procurement and manufacturing far enough in advance that once they receive final approval, they’re ready to start putting steel in the water.
The Mod Rule says that, so long as your third-party certified verification agent is involved, developers are free to conduct onshore fabrication while the permitting process is going on.
In what ways does the ruling streamline overly burdensome processes?
Beyond what I’ve already mentioned, there are three key examples of how the rule streamlines the permitting process:
First, it gets rid of the site assessment plan requirement for deployment of meteorological (met) buoys.
The original rule assumed that developers were going to be installing met towers that are pile driven into the seabed as tall as the hub height of a wind turbine. Installation of these structures could have environmental impacts, and so it made sense to have a separate plan for that.
But it quickly became apparent that developers were shifting rapidly to the use of met buoys for measuring wind speed. A buoy is a buoy, and they are routinely deployed off our coasts with minimal permitting because they have minimal environmental effects. But under the original BOEM regulations, developers were waiting a year or more to be allowed to deploy their met buoys. Now, developers just need to get a general permit from the U.S. Army Corps, which approves buoys within a matter of weeks using a standard set of conditions. That is all time and money saved by developers, with knock-on benefits from being able to start collecting wind speed data much sooner.
Second is the timing of data submittal. The original regulations required that your COP had to have a geotechnical exploration at every single turbine location. But with the PDE approach, you may not know exactly where your turbines will be located at the time of COP submittal.
BOEM went in the direction of offering a more flexible and performance-based approach to data submittal. You still must present enough data for engineering or environmental purposes, but the data can be provided when it’s needed. The geotech at every turbine location can now be submitted after COP approval when BSEE is reviewing the final engineering reports. A third example of streamlining is BOEM’s reform of the offshore wind lease structure. Under the original regulations, the operations term started when you got your COP approval. The default length was only 25 years, which BOEM realized was too short of time, given the ever-expanding design life of these devices. But it was only 22-23 years because it can take several years after COP approval to construct your project and put it in commercial operation.
Now, the operations term starts once you’re completely done and are generating energy commercially on all your wind turbines, and it lasts 35 years instead of 25. That’s much more certainty.
Will this ruling help with offshore transmission issues?
This is an area where the rule could help in a small but significant way.
The U.S. offshore wind industry is eventually going to need to move from radial transmission where every single project has its own line to shore to some sort of shared offshore grid. One of the challenges has been: How do you align the process of siting shared transmission, which would collect from multiple projects, with the process of siting each individual wind farm? How do you make sure that the shared transmission is going to be ready in time to allow individual projects to plug in so they’re not just sitting there waiting for transmission?
BOEM is required by statute to determine if there is competitive interest in offshore transmission rights-of-way before they are issued. In the Mod Rule, BOEM signaled that, if a company has won a state request for proposal (RFP) for stand-alone offshore transmission, that could be a major factor in determining whether there is competitive interest in the property right needed to build that transmission line offshore. In effect, BOEM is saying, “States, you go first. You do RFPs to figure out who you want to build these transmission lines, find the most qualified companies from a technical and economic perspective, and we will run our competitive process in a way that heavily factors in whether a transmission developer has secured a state RFP.” This approach eliminates a key source of uncertainty, and thus helps facilitate this future world where we have backbone transmission.
What challenges might companies and stakeholders have to face in light of the ruling?
Because much of the rule involves codifying practices that were already being done on a case-by-case basis, I actually don’t think the transition for developers is going to be that hard.
Stakeholders are generally not affected by the rule. The Mod Rule is really a model of good government. BOEM and BSEE have improved their processes in rational, well-considered ways, but the compromising the rigor of their environmental and safety analyses.
What do you feel is the industry’s reaction to the new rule?
I think it is generally positive. BOEM and BSEE hit a solid double here, but I think they maybe missed some opportunities that would have made it a home run.
It would have been very helpful, for instance, if BOEM had bound itself to a more predictable permit review timeline. NEPA does require you to go through your NEPA process from a notice of intent to a final environmental impact statement within two years with some exceptions, but there’s less certainty in terms of the length of time between submitting your COP and starting the NEPA process on the front end, and then going from your final environmental impact statement to your approval.
We could have also used more certainty regarding leasing. The rule includes a requirement that BOEM issue a leasing plan every few years, but it would be helpful if they outlined what factors they will be using to ensure a future administration is constrained from just doing no additional leasing. This is critical because the U.S. industry needs a pipeline of new projects to help build and sustain a domestic supply chain and ensure that states can meet their ambitious offshore wind energy mandates and goals.
Berge Bulk’s Berge Neblina, a 388,000 dwt Valemax Ore Carrier, is completing its voyage to Brazil following the successful installation of four 5x35m rotor sails from Anemoi Marine Technologies Ltd.
The installation, which took place during the vessel’s scheduled dry docking, was completed at Yiu Lian Dockyards (Shekou) Ltd in China. The selected rotor sails have been installed on Anemoi’s folding deployment system, where the sails can be folded from the vertical to mitigate impact on air draught and cargo handling operations when in port.
“Leveraging the latest in wind technology to reduce our fleet’s emissions is an important part of Berge Bulk’s ‘Maritime Marshall Plan’ for decarbonization,” said Paolo Tonon, Berge Bulk’s technical director. “We are optimistic that these rotor sails can deliver up to 8 percent carbon reduction.”
“Anemoi’s collaboration with Berge Bulk demonstrates how we are both working in partnership to ultimately secure shipping’s zero-emission future,” said Kim Diederichsen, Anemoi CEO. “Anemoi remains committed to maintaining its position as a leading provider of critical vessel decarbonization technology.”
Rotor sails, also referred to as Flettner Rotors, are comprised of vertical cylinders that, when driven to rotate, harness the renewable power of the wind to propel ships. These mechanical sails capitalize on the aerodynamic phenomenon known as the Magnus Effect to provide additional thrust to vessels. By leveraging wind energy, Berge Neblina will see increased efficiency by reducing the load on the main engine while maintaining speed, reducing fuel consumption and resulting in fewer greenhouse gas emissions.
The technology is being increasingly embraced by ship owners, especially in the bulk sector, who are aiming to achieve net-zero shipping emissions. Rotor sails have emerged as a preferred technology to augment and enhance the energy performance of vessels. Rotor Sails are a compact technology, which offer a large thrust force to propel ships, helping them comply with pivotal international emission reduction benchmarks such as the Carbon Intensity Indicator (CII) and EEDI/EEXI.
North Star, specialist vessel operator for offshore infrastructure support services, was recently announced as the first mover on the Midi-SOV — a new offshore wind ship design developed by Chartwell Marine, a pioneer of next-generation vessel design, and VARD, a leading designer and shipbuilder of specialized vessels.
The Midi-SOV is a 55-meter offshore wind craft ready for build in the European, Asian, and U.S. markets. North Star has entered an agreement with Chartwell and VARD, becoming the first to adopt and use the Midi-SOV on offshore wind projects, investing in upfront design fees to facilitate vessel construction for European operations.
“We designed the Midi-SOV with a clear vision of its integration into future offshore wind fleets, filling the gap that had emerged between CTVs and SOVs and addressing key operational challenges as the wind industry evolves,” said Andy Page, Chartwell Marine managing director.
“Together with VARD, we have been very encouraged by the positive response we’ve received from offshore wind operators, underscoring the industry’s readiness for new solutions that enhance efficiency, safety, and overall project costs. And, of course, we are delighted to continue our collaborative relationship with North Star as they take a leading role in bringing this vessel from design to reality,” Page said.
Chartwell and VARD’s Midi-SOV solution addresses challenges in the offshore wind sector by bridging the gap between Crew Transfer Vessels (CTVs) and Service Operation Vessels (SOVs), providing comfort and workability while offering a cost-effective alternative to full scale SOVs. With a design based on operational data to meet the niche requirements of offshore wind developers and operators, the Midi-SOV is intended to complement existing fleets.
The operational efficiency of the Midi-SOV was identified as one of its key advantages, evidenced by lower technician attrition rates due to the comfortable and spacious working environments provided. Furthermore, discussion included the Midi-SOV’s robust safety performance, particularly in reducing risks during technician transfers and crane operations.
“We’re excited about the operational versatility the design can give us, as well as the high standards of safety, availability and cost efficiency it promises — and proud to play our part in bringing the first Midi-SOVs to market,” said Andrew Duncan, North Star’s renewables and innovations director.
Clearway Energy Group closed financing and has begun repowering construction of its Cedro Hill wind farm in Webb County, Texas. The repower will increase Cedro Hill’s capacity to 160 MW from 150 MW. Once complete, the repowered project will generate enough electricity to power more than 40,000 homes during peak hours.
“Our Cedro Hill repowering is our most recent example of how Clearway is upgrading its sizeable existing fleet to deploy resilient, state-of-the-art technology on sites that have proven, high-quality wind resources,” said Chris Fox, senior vice president of Construction at Clearway Energy Group. “As a long-term owner and operator, we are pleased that repowerings like Cedro Hill deliver decades of more value for our local landowners and communities.”
Clearway Energy is repowering its Webb County, Texas, wind farm. (Courtesy: Clearway Energy Group)
This project will mark Clearway’s fifth wind-farm repower in Texas and sixth across its portfolio, amounting to more than 700 MW of upgrades to deploy resilient technology on sites with strong wind resources.
The Cedro Hill repower represents a $269 million investment in south Texas. Upon completion, the repower will add another 15 years to the project’s operating life and extend property taxes and landowner payments to Webb County by $27 million.
The repower will replace the blades and nacelle with General Electric (GE) equipment across the site’s 100 turbines, and manage all replaced equipment. In collaboration with Wanzek Construction, Inc., a MasTec Renewables company, Clearway is undertaking efforts to ensure that existing materials, including fiberglass, are recycled or diverted from landfills. More than 200 skilled laborers will support construction.
Cedro Hill was built and commissioned in 2010, with its generated power sold under a long-term power purchase agreement with CPS Energy, the nation’s largest municipal electric and gas utility, serving the city of San Antonio, Texas, and one of the nation’s largest municipal buyers of wind energy. As part of the repowering, CPS Energy extended its existing agreement to support its commitment to growing its renewable energy portfolio. CPS Energy continues to benefit from 100 percent of the power generated by the Cedro Hill wind farm.
“MasTec Renewables is looking forward to another successful project with Clearway Energy Group,” said Brendon Lamppa, director of construction, MasTec Renewables. “This will be the third overall, but the first repower project that the teams will have worked together on. This is a unique opportunity given the fact that MasTec Renewables’ legacy personnel performed the original build in 2010. The repower aspect also brings a recycling component into the equation. All 100 existing hubs and 300 existing blades will be removed, cut on-site, and shipped to a recycling facility to be processed for beneficial reuse.”
The Bureau of Ocean Energy Management (BOEM) recently announced its approval of the New England Wind Construction and Operations Plan (COP), which authorizes construction and operation of two wind-energy projects.
This is the final approval of these two projects from BOEM, following the agency’s April 2024 Record of Decision.
The New England Wind projects are about 20 nautical miles south of Martha’s Vineyard. (Courtesy: BOEM)
“The Biden-Harris administration is committed to advancing offshore wind-energy projects like New England Wind to create jobs, drive economic growth, and cut harmful climate pollution,” said BOEM Director Elizabeth Klein.
“We are proud to announce BOEM’s final approval of the New England Wind projects. They represent a major milestone in our efforts to expand clean-energy production and combat climate change.”
The approval will permit the construction and operation of two offshore wind-energy facilities, known as New England Wind 1 and New England Wind 2, which together will have a total capacity of up to 2,600 MW of renewable energy that could power more than 900,000 homes each year.
The two projects are situated about 20 nautical miles south of Martha’s Vineyard, Massachusetts, and about 24 nm southwest of Nantucket, Massachusetts. The COP for the two projects includes up to 129 wind-turbine generators, up to five electric service platforms, and up to five offshore export cables transmitting electricity to onshore transmission systems in the Town of Barnstable and Bristol County, Massachusetts.
BOEM considered feedback from Tribes, other government agencies, ocean users, and others prior to the decision.
The feedback resulted in required measures to avoid, minimize, or mitigate any potential impacts from the project on marine life and other important ocean uses, such as fishing.
Since the start of the Biden administration, the Department of the Interior has approved eight commercial-scale offshore wind-energy projects in federal waters, and BOEM has held four offshore wind lease sales, including offshore New York, New Jersey, the Carolinas, and the first-ever sales offshore the Pacific and Gulf of Mexico coasts.
Modulift, a lifting equipment manufacturer, is facilitating the construction of a complex offshore energy project in the North Sea. Deemed to be the world’s first artificial energy island, “Princess Elisabeth” marks a milestone in renewable energy infrastructure development.
Situated 30 miles off the coast of Belgium, in the 3.5-GW Princess Elisabeth offshore wind zone, the island will act as an international energy hub to centralize all electricity produced by wind farms in the zone.
Construction on the Princess Elisabeth island will run until the end of 2026. (Courtesy: Modulift)
TM Edison, a joint venture between marine companies Jan De Nul and DEME Group NV, is responsible for the construction and installation of the energy island, which will contribute to the EU’s goal of 300 GW offshore wind capacity by 2050.
Assembling the infrastructure of the energy island requires lifting and maneuvering of large structural elements and equipment. The outer perimeter of the island will be made up of a series of concrete structures known as caissons. These are built onshore before being transported to their offshore location.Construction of the Princess Elisabeth Island is expected to last until the end of 2026.
“Modulift’s spreader beams have been integral to the construction of the concrete sections, with one formwork section weighing around 17 tons and 10 meters in length. We are using a 1-over-1 configuration, utilizing a MOD 34 spreader beam at the top and a MOD 24 on the bottom; Modulift spreader beams have been instrumental in safely hoisting the formwork into position,” said Ruben Verschueren, TM Edison’s site superintendent, civil works.
“Modulift is known for its iconic yellow spreader beams across the renewable and offshore energy sectors, and has set the precedent for safety, efficiency, and fast delivery in the global lifting industry,” said Sarah Spivey, managing director.
Wind turbines play an increasingly important role in the global energy market. The World Wind Energy Association [1] reported yearly growth of 13 percent in 2021 to an overall capacity of 840 GW, enough to provide more than 7 percent of the global power requirement. Operating requirements are becoming more demanding with the move toward larger structures, higher loads, increased off-shore development and hostile environmental conditions. Downtime is expensive both in lost energy production and repair costs, particularly for off-shore turbines. Wind turbines are particularly prone to gearbox failure and a recent paper [2] reported more than 50 percent of this is due to bearing problems. Rolling bearings in wind turbines are in the gearbox, shaft, pitch/blade, yaw, and generator systems, where they are often subject to extreme operating conditions of high loads, low temperatures, and variable wind speeds [2]. Various lubrication-related failure modes have been identified, including scuffing, micro-pitting, and fretting corrosion [2], [3].
Metal gear in oil. Car service concept. Oil change in mechanisms, engines, motors.
1 Wind turbine tribology
Both grease and oil lubricants are used in wind-turbine systems; grease is the main lubricant in the pitch/yaw, generator, and shaft bearings, and formulated oil is in the gearbox. One of the challenges is to optimize grease formulation for such demanding conditions, particularly low-temperature operation.
The focus of this article is the low-temperature lubrication of the pitch and yaw bearings, as they can fail due to fretting/corrosion and false brinelling [2] damage. Fretting is defined as “small amplitude relative oscillatory motion between two contacting bodies that gives rise to both wear and fatigue” [3]. It is associated both with low amplitude motion and system vibration. The origins of oscillation and fretting in bearings are due to the operating requirements of the turbine.
While most bearings rotate, oscillation is not uncommon. In some applications, the oscillation is intended and fairly large. The pitching system for wind-turbine blades is an example. The blade angle is changed to accommodate changes in wind speed and keep the rotating speed of the propeller within acceptable limits. In older, smaller wind turbines, the angle was changed based on the overall wind speed, and the changes were not very frequent. In the more modern, larger wind turbines, the blade angle is often changed during the rotation. This helps to compensate for differences in wind speed at different heights during the rotation and is known as active pitching [3].
Figure 1: HFRR test device: fretting condition, simple schematic and installed in the freezer.
There are many examples of unintended oscillation. In wind-turbine blade bearings, for example, at the top of the rotation, the blade is pushing down, axially on the bearing. At the bottom of the rotation, it is pulling the bearing axially in the opposite direction, which leads to a small axial movement. Other oscillations can be caused by wind-speed changes (gusting) and offshore wave motion.
The bearings in the blade pitch system are usually four- or eight-point contact rolling bearings [2]. Stammler et al. [3] [4] provided an analysis of pitch angle rotation for a 7.5-MW turbine blade bearing; this was a 4,690 mm diameter double row bearing with 80 mm diameter balls. The pitch oscillations are low frequency and typical pitch angle of a few degrees (<2°). The analysis concluded that 15.3 percent of all pitch movements occur between angles of 0.03° and 0.2° [4], which correspond to fretting conditions.
Fretting occurs when the reciprocating motion (L) is less than the Hertzian contact width (2a) [3], [4], [5] (Figure 1), which has implications both for the formation and behaviour of wear particles [5] and the ability of the lubricant to generate a separating film. The restricted motion means wear particles generated during rubbing are not expelled from the contact zone and remain actively involved in the lubrication process [5]. A second outcome is reduced entrainment of lubricant impeding the development of a sufficient lubricating film as the restricted motion and low speed combine to hinder the formation of an elastohydrodynamic lubricating (EHL) film [6].
Greases used in wind turbine bearings typically use calcium or lithium-based thickeners with either mineral oil or synthetic hydrocarbon base oils. Several standard tests are used to assess grease fretting performance. ASTM D4170 [7], known as the Fafnir fretting test, measures wear in thrust bearings run at a frequency of 30 Hz, 12°arc of motion and room temperature. Bearing wear is measured by weight loss. A second test ASTM D7594 [8] uses a reciprocating ball-on-flat bench test (SRV) to measure friction and wear typically at 50 °C, 0.3 mm stroke, 50 Hz and 2.2 GPa. Neither of these tests is usually run at low temperatures representative of wind turbine operation. The NLGI have published a specification guide (HPM+LT) for greases used at low temperatures [9]. The tests include ASTM D1478 which is a low temperature (−30 °C) ball bearing torque test and other flow/grease mobility tests [9]. A low temperature fretting test appropriate for wind turbine applications is not specified. At present none of the industry tests accurately reproduces the fretting conditions in low-temperature wind turbines.
Figure 2: Wear measurement using SWLI.
A number of papers have reported grease lubrication under fretting conditions both in bearings and ball-on-disc tests [10], [11], [12]. The Schwack et al. study [10] focused on conditions in wind turbine pitch blade bearings and compared measurements in a bearing and SRV bench test for six commercial greases. The greases had a range of thickener types (calcium/calcium complex, lithium/lithium complex) and base oil viscosities/composition. The bearing tests were run at fretting (L/2a=0.9) and sliding conditions (L2a=13.3, 29.2). The SRV tests were run at L/2a = 1 at 8 and 32 Hz, stroke length 0.3 mm and 20 °C. The overall conclusion was that greases with low base oil viscosity and high bleed rates gave the best fretting wear performance which is in line with other studies [11][12].
Despite the importance of wind turbine technology and the increasing power demands, currently we do not have a simple bench-test which will replicate wind turbine bearing fretting at different temperatures. The aim of the current work was to address this problem through the development of a bench-top fretting test which could be run at very low temperatures, down to minus-40°C. The tests were designed to replicate low amplitude fretting in bearings under realistic operating conditions.
A series of ball-on-disc reciprocating fretting tests was carried out to investigate the friction and wear properties of four greases at different temperatures (25, 7, minus-20, minus-40°C). Post-test the wear scar and grease distribution on the disc was examined by a low-power microscope. The wear scar dimensions on the ball were determined by White Light Interferometer (WLI). Further analysis was carried out to understand grease lubrication mechanisms under fretting conditions. The chemical composition of the grease film in the wear scar was analysed by Reflection Absorption Fourier Transform Infrared Spectroscopy (RA-IRS). The disc was then cleaned, and Raman Spectroscopy and SEM-EDS used to examine the surface chemistry of the wear scar.
Figure 3: Example friction coefficient and ECR traces for greases A (upper) and D (lower) at 25°C.
2 Test methods and materials
2.1 Friction and wear tests
The High Frequency Reciprocating Rig (HFRR) [PCS Instruments, Acton, UK] was initially designed for evaluating diesel fuel lubricity as described in ASTM D6079 and ISO 12156 [13], [14]. It has also been extensively used to evaluate wear and friction properties of a wide range of lubricants, additives and greases. The HFRR set-up is schematically shown below in Figure 1, in which a steel ball (6 mm diameter) is loaded and reciprocated against a stationary steel disc (10 mm diameter).
The surface roughness (Ra) of the ball and disc specimens was 9.9 ± 1.2 nm and 4.8 ± 0.8 nm respectively; while the hardness of the ball and disc specimens was 800 Hv and 196 Hv respectively. The ball is driven by an electromagnetic vibrator to achieve oscillating motion at desired frequency and stroke length. A temperature probe is under the disc specimens to monitor the test temperature. A force transducer mounted underneath the heater block measures real-time friction force throughout the test. Before testing, metal specimens were cleaned in an ultrasonic bath with toluene, rinsed in acetone, and air-dried. At least three tests per condition were run.
The formation of a tribofilm or any change to the contact between the ball and disc specimens was indicated by the electrical contact resistance (ECR). The ball-disc contact resistance is related to ball-disc contact geometry, resistance of lubricant, and/or specimen materials, and formation of a tribofilm. During a HFRR test, a constant electrical potential of approximately 15 mV is applied to the ball and disc contact with a balance resistor connected in series. Therefore, when the ball and disc are fully separated (open circuit), the voltage will be c.a. 15 mV, representing 100 percent film; while, when there is direct metal-to-metal contact between the ball and disc, the voltage will be 0 V, representing zero percent film. Although ECR does not provide a direct measurement of film thickness in association with friction and wear results, it can provide insights into possible lubrication and wear mechanisms.
Table 1: HFRR test conditions.
To replicate grease lubrication conditions occurring in wind turbine bearings it is necessary to control a number of parameters, these include stroke length for fretting, temperature, and initial grease sample application.
Liquid lubricant tribology tests are usually run with an excess of fluid present. However, in bearings, there is a very small amount of grease situated close to the contact and lubricant replenishment of the raceway is often limited [16], [17]. For most liquid lubricant tests, it is sufficient to replicate the speeds, temperature, contact pressure, etc. For greases, it is also necessary to consider the mechanisms driving lubricant replenishment of the rubbed track. Clearly, this is only really achieved in full-scale bearing tests; the aim of the HFRR test design was to replicate the relevant conditions (motion/temperature/pressure), and lubricant distribution around the contact. To replicate this condition in the HFRR test, the grease sample was applied as a thin film on the disc. If an excess of grease is used, it can be dragged back into the contact zone by the ball holder, rather than the ball motion. By using a controlled, thin film sample, the aim was to test the ability of the grease to be retained in the contact or replenish the rubbed track after it is displaced during fretting. To meet this requirement, the grease was applied as a circular spot (4 mm diameter and 50 µm thick) on the disc.
The HFRR test conditions are listed in Table 1.
Figure 4: Example friction coefficient and ECR traces for greases A (upper) and D (lower) at minus-40°C.
2.2 Post-test analysis
2.2.1 Wear analysis
A Bruker Scanning White Light Interferometer (SWLI) Contour GT was used to quantify wear scar diameter (WSD) on the ball after the HFRR tests. It uses scanning white light interferometry to map surface topography to sub-nanometer resolution. An example of ball WSD measurement is shown in Figure 2. The SWLI scans the topography of ball surface sector, which includes the wear scar, and the analyzing software fits the curvature of the surface using the ball radius of 3 mm. The fitted surface is thus a plane as shown in the top picture in Figure 2. Green and red cursers are then located on the edge of the wear scar to measure the lengths between them. The WSD was the average horizontal and vertical lengths of the ball wear scar. In Figure 2 the horizontal and vertical lengths of the ball wear scar are 195.9 µm and 147.1 µm so that the average WSD is 171.5 µm.
2.2.2 Grease composition analysis:
FTIR and RA-IRS spectroscopy
Fourier Transform Infrared Spectroscopy (FTIR) is used to characterize and identify chemical species particularly for organic compounds. In this study, it was used to detect thickener and base oil components of the greases [16]. A PerkinElmer Frontier FTIR equipped with an IR microscope was used. Fresh greases were also analyzed and used as reference spectra for the thickener/base oil content. Post-test grease films remaining in the disc wear track were examined using the IR microscope. The sample is first observed using the visual function and the sample area defined by a 100 µm diameter aperture. The microscope is then switched to the IR mode and the sample area scanned using the reflection-absorption method RA-IRS (Reflection-Absorption Infrared Spectroscopy). The in-track spectra were compared to the fresh grease reference to identify any changes in composition due to rubbing.
2.2.3 Analysis of wear scar surface chemistry: Raman spectroscopy and SEM-EDS
Raman Spectroscopy was used to identify chemical species in both organic and inorganic compounds. An Alpha 300 RA (WiTec, Germany) with a 532 nm laser source was used to analyze the disc wear tracks after the HFRR tests, to investigate the formation of iron oxide films [18]. The samples were solvent-cleaned and spectra taken from different points in the wear scar. Representative results are shown in the paper for this analysis.
Table 2: Grease test samples: thickener types and base oil viscosity.
Raman spectroscopy was used to identify iron oxides in the wear track; however, additive tribofilms formed in the disc wear tracks were not identified. Scanning Electron Microscopy with Energy Dispersive X-ray Spectroscopy (SEM-EDS) was therefore used to investigate the element composition of the disc wear tracks. SEM enables a high-resolution image of the wear track, thus the same area of the wear track, which were studied with Raman, could also be analysed by EDS to determine the chemical composition. A Tescan Mira SEM-EDS (Kohoutovic, Czech Republic) was used to take high magnification and resolution images and detect elemental composition in the HFRR disc wear tracks.
2.3 Materials
Four greases were investigated, details of which are in Table 2. The greases chosen were commercially available from different manufacturers and representative of formulations currently used in wind turbines. All the greases were hydroxstearate-based with calcium (A, B) or lithium/lithium complex (B, C, D) thickener chemistries. As these were commercial samples, it was not possible to obtain the base oils for viscosity measurements, so the base oil viscosities at minus-20 and minus-40°C quoted in Table 2 were calculated using the ASTM viscosity-temperature chart based on the published viscosities measured at other temperatures.
Figure 5: Averaged friction (upper) and ECR (lower) results.
3 Results and discussion
3.1 HFRR friction and wear results
Friction and ECR values were recorded throughout the test and are typically plotted against time or number of strokes. At least three tests were carried out for each grease/temperature combination. Overall, the results were very repeatable; examples are shown in Figure 3, Figure 4 for greases A and D at 25 and minus-40°C, as these represent the performance extremes.
At the higher temperatures (25 and 7°C) the friction coefficient was low and stable for A, B, and C (<0.25). The exception was grease D, where the CoF was initially high (µ = 0.5-0.8) and unstable; it then declined suddenly to a steady value (µ∼0.15) for most of the test. Grease A ECR value increased steadily after 40,000 strokes (see Figure 3 upper graph) to values of 20-60 percent, which might indicate the development of a reacted additive film. For the lower temperatures (minus-20, minus-40°C), friction traces became increasingly high and unstable. Examples are shown in Figure 4 for A and D at minus-40°C. At minus-40°C, greases B, C, and D had very high, unstable friction and ECR traces. The exception was grease A, where the ECR trace was zero.
Figure 6: Summary of ball wear scar diameter measurements.
The test average friction coefficient and ECR results are plotted in Figure 5. The friction chart shows that, at higher temperatures (25 and 7°C), the four test greases presented comparable friction coefficient (µ ∼0.15), with slightly higher values for grease D (µ > 0.2). At minus-20°C, greases A, B, and C gave mixed results, the CoF was often unstable and varied in the range µ = 0.18–0.6. Grease D gave very high (µ = 0.6–0.8), unstable friction traces in all tests. At minus-40°C, all greases gave very high average friction coefficients (0.5–0.8).
Figure 5 (lower graph) shows the average ECR value of three tests. At 25 and 7°C, greases A and B recorded a higher ECR, which gradually increased after ∼30,000 strokes to levels of 30-60 percent and 15 percent, respectively. Greases C and D recorded much lower ECR (<2%) at both temperatures. At the lower temperatures minus-20 and minus-40°C), greases A and B recorded very low ECR values (<5%). The ECR traces for greases C and D were very high and unstable.
The ECR measurement indicates a non-conductive material is formed in the interface. The ball-disc contact resistance varies with contact geometry, rubbing materials, tribofilm, and resistance of lubricant and wear particles. Therefore, if the ECR is low, it represents there is predominantly direct surface contact. However, if the ECR is high, it may indicate there is a high electrical resistance tribofilm [19] or wear particles formed in the contact [20]. These two cases will lead to either low or high wear respectively, because a tribofilm normally protects rubbing surfaces, but debris particles could accelerate abrasive wear. Further wear scar chemical analysis will help to identify whether it is a tribofilm or wear debris.
Figure 7: Images of the disc wear scar and grease films for A and D tests at 25°C (upper) and minus-40°C (lower). Scale marker: 2,000 µm.
The ball wear scar measurements are summarized in Figure 6. All the greases gave similar results at 25 and 7°C; however, at minus-20°C, grease D wear scar was significantly larger. At minus-40°C, greases C and D gave increased wear levels compared to A and B. The high wear recorded for greases C and D at the lowest temperatures corresponds to the high ECR levels recorded. This would suggest the ECR response is due to the formation of non-conductive wear debris [20] rather than additive film formation.
3.2 Microscope Images of Wear Scars
At the end of the test, the disc was not cleaned but retained with the grease film for further analysis. Initial inspection was by a low-powered microscope and examples are shown in Figure 7 for samples A and D tested at 25 and minus-40°C.
For most tests, grease is retained close to the wear scar where a thinner film is seen within the reciprocating zone. Outside this zone, the grease is undisturbed. The image for grease D at minus-40°C was strikingly different. The wear scar is seen clearly as the grease film appears to have retracted from the sheared region. Again, outside this zone, the grease is undisturbed.
Larger-scale magnification images of the A and D wear scars are shown later in the article with the Raman analysis (Figure 10). For grease D, the characteristic red-brown wear debris indicating haematite (Fe2O3) is clearly visible around the wear scar. This is indicative of fretting corrosion damage [5], [10], [12] and is probably the origin of the high ECR value measured for this test [20].
Figure 8: FTIR spectra of fresh greases.
3.3 FTIR Results
The fresh grease spectra are shown in Figure 8 for a limited wavelength range of 1,800-1,000 cm-1, as this region contains the most important peaks used to identify thickener, base oil, and additive grease components [16]. The spectra are plotted in Absorbance units, after normalization of the C-H peaks at ∼1,460 cm-1, as this allows for comparison of the thickener content.
The spectra of fresh greases were saved as reference for comparison with RA-IRS spectra of lubricant films in the wear scar. The strongest absorbance peaks are associated with the thickener and base oil components, while those due to additives are usually weak. Peaks at 1,580 and 1,560 cm-1 correspond to the asymmetric stretch vibrations of the carboxyl group (COO) in the hydroxystearate thickener [21]. The bands at ∼1,464 and 1,377 cm-1 are various CH vibrations due to both base oil and thickener. From this relative comparison, the thickener content is approximately D>C>B>A. Any change to the composition of the grease film can be seen as a relative change in the component peak absorbance or loss compared to the spectrum of fresh grease [16], [21].
Figure 9: Upper: Microscope images for grease samples for A and D at minus-40°C indicating the area of RA-IRS analysis. Scale marker 250 µm: Lower: FTIR spectra of fresh and wear scar grease samples for A and D at 25 and minus-40°C. Grease A at minus-40°C. Grease D at minus-40 °C.
Post-test lubricant deposits in and around the disc wear track were examined by RA-IRS, and the results were compared to the reference fresh grease spectra. Figure 9 (upper) shows images from grease films after tests at minus-40°C; typical RA-IRS sample positions are indicated by red circles. The corresponding RA-IRS spectra are shown in Figure 9 (lower).
For all greases at higher temperatures (25 and 7°C), both thickener and base oil were present in the wear scar. The characteristic bands corresponding to thickener at 1,580 and 1,560 cm-1 can be seen in Figure 9 (upper) at 25°C for both A and D. Similar results were recorded for the 7°C tests. However, at lower temperatures, complete loss of the thickener peaks was seen for grease D at minus-20°C. Thickener was still present in the wear scar films for greases A, B, and C. At minus-40°C, thickener was present for grease A but not for greases C and D (Figure 9 lower). The RA-IRS results for grease B at minus-40°C were ambiguous, although the grease gave relatively low wear (average WSD 250 µm), thickener was not identified in the wear scar. However, the spectra were poor quality, and clear interpretation is not possible due to baseline drift. Interestingly, although the thickener is absent for some greases at low temperatures base oil peaks (1,500-1,300 cm-1) are still present, suggesting released oil reflow into the track is still occurring. The implications of these findings will be discussed in Section 3.5.
Figure 10: Raman and SEM-EDS analysis on disc wear tracks from tests on from 25 °C tests: upper grease A and lower grease D. Red circle shows analysis position. Scale marker: 60 µm.
3.4 Raman Spectroscopy and SEM-EDS Analysis of Wear Scar Chemistry
Raman spectroscopy and SEM-EDS were used to investigate the surface chemistry of the disc wear tracks. After each HFRR test following the RA-IRS analysis, the discs were rinsed with toluene and air-dried. Any grease remaining in the wear tracks was washed away, so the Raman and EDS analysis was focused on the adherent tribofilm and wear debris. EDS analysis was carried out for different regions and representative results are shown. The results are presented as element composition in atomic percentage. Figure 10 and Figure 11 show example Raman spectra and SEM-EDS analysis for greases A and D at 25 and minus-40°C.
The SEM images provide evidence of different wear mechanisms. For grease A at 25°C (Figure 10 upper), the surface is grooved, which is characteristic of a predominately abrasive wear mechanism. The friction coefficient is low throughout the test (µ∼0.15), while the ECR trace increases to 20-40 percent as the test proceeds. EDS results show some phosphorus and sulphur (0.01 at%), which might indicate formation of a non-conducting P/S additive film [19]. There were also high percentages of iron and oxygen confirming the iron oxides detected by Raman. It is worth noting the working depth of EDS is up to 5 µm depending on the accelerating voltage, but Raman scanning depth is from nanometers to 1 µm; while the thickness of anti-wear tribofilms are in nanometers. Therefore, it is reasonable that the percentage of iron was high, but sulphur and phosphorus content were low; in addition, chromium and carbon (resulted from oxidation/degradation of lubricants) were detected.
Figure 11: Raman and SEM-EDS analysis on disc wear tracks from tests on from − 40 °C tests: upper grease A and lower grease D. Red circle shows analysis position. Scale marker: 60 µm.
In contrast, point 2 did not register sulphur and phosphorus. Instead, the relatively higher oxygen content clearly coincided with Raman results, suggesting the formation and accumulation of iron-oxide debris. The low wear (178.9 µm average ball WSD) is possibly due to the combination of additive film formation and the presence of thickener in the lubricant film as indicated by the RA-IRS results (Figure 9 lower).
In contrast, the wear scar for grease D at 25°C (Figure 10 lower) appears “plucked” with smeared regions, indicating adhesive wear and plastic deformation. The EDS analysis indicates high sulphur content (0.04 at%) in the wear scar, which is supported by peaks in the Raman spectra at 342/374 cm-1 tentatively assigned to an FeS film [22]. This analysis is further supported by the friction and ECR results. The friction coefficient is initially high and erratic (µ = 0.5–0.7), dropping to a low stable value ∼µ = 0.18 after 20,000 strokes. The ECR value was low throughout most of the test. These observations, coupled with the relatively low wear value (247.2 µm average ball WSD), indicate the formation of an electrically conductive sulphur-containing additive film [22] as the test progressed.
The results for greases A and D at minus-40°C are shown in Figure 11. Wear remained low for greases A and B (∼250 µm) but increased significantly for greases C and D (400 and 500 µm respectively). Friction also increased for all greases, for A, C, and D, average friction coefficient was 0.5-0.8. Grease B was lower (∼µ = 0.3), but there was a large variation in results from the individual tests (µ = 0.2-0.9).
For grease A, there appear to be compacted deposits in the wear track (Point 1 Figure 11 upper image). The EDS analysis indicates high carbon and fairly low oxygen content. The Raman spectra identify Fe3O4 (peak at 666 cm-1) and possibly γ − Fe2O3 (1312 cm-1). There is also an indication of organic content, possibly due to Ca thickener (CH ∼ 2,427 and C O _ 1,570 cm-1). This material has survived solvent cleaning and is tightly bound, possibly part of a transformed layer of wear debris and Ca thickener. The RA-IRS spectra also indicated the continued presence of thickener in the lubricant film. As such, it contributes to the low wear but relatively high friction measured for grease A. S/P films were not detected in the wear track.
Table 3: Summary of Raman peak results.
The SEM image from grease D also shows deposits in the wear scar, although this appears to be more particulate. Low magnification images of the wear scar show red iron oxide deposits (Figure 9 lower), and this is supported by the Raman and EDS analysis. Fe2O3 (doublet 220, 281 cm-1) was detected by Raman spectroscopy. In addition, high levels of oxygen (45-50 at%) were measured by EDS. The test also recorded very high and unstable ECR trace, which is due to non-conductive iron oxides formed during fretting wear [20] (Table 3).
3.5 Discussion: lubrication mechanisms and implications for blade bearing grease formulation
The article reports friction and wear results for a range of commercial greases. Although the results are limited, it is possible to infer some relationship to grease composition. The greases had various calcium or lithium hydroxystearate/complex thickener systems at different concentrations. Some information on the thickener concentration can be gleaned from FTIR spectra of the fresh greases (Figure 8). The peak absorbance for the thickener (C_O ∼1,580 cm-1) relative to the base oil (C-H ∼1,460 cm-1) can be used to give a rough guide to thickener content.
The spectra were normalized to the CH peak absorbance so a direct comparison is possible and this indicates the thickener concentrations were ranked D>C>B>A. Similarly, for the base oil, there is limited information for the composition. The specified base oil viscosities varied from D (1,116 mm2/s)>C (316 mm2/s)>B (224 mm2/s)>A (23.3 mm2/s) at 25°C. It was possible to extrapolate base oil viscosities from published data for all the test temperatures using ASTM plots and this is summarized in Table 2. The FTIR analysis suggests that greases A and B have a synthetic base oil with an ester component (peak at 1,735 cm-1). It is assumed all greases contained a range of additives and evidence of P- and S-containing reacted films was obtained in the EDS analysis.
Differences were recorded in the measured wear at the lower temperatures, and this can be related to the grease composition and mechanism of film formation. Wear performance was better for the lower viscosity base oil and lower thickener-content greases. The relationship of measured wear generally increases with base oil viscosity as shown in Figure 12, which is the inverse of the usual EHL film thickness “rules.”
Figure 12: Wear scar diameter plotted against base oil viscosity for all greases and temperatures.
Further evidence for the lubrication mechanism comes from the RA-IRS analysis of the film remaining in the wear scar at the end of the tests. At higher temperatures (25, 7°C), thickener was present in lubricant film for all greases, and correspondingly similar low friction and wear results were recorded. At the lower temperatures, thickener peaks were absent for greases C (minus-40°C) and D (minus-20, minus-40°C), although a thin film of free oil was present, and this condition was associated with much higher wear. The anomaly in this analysis was grease B, which had low wear at minus-40°C but with no apparent thickener present. However, the analysis also showed the presence of an ester component in the lubricant film, which might contribute to enhanced friction and wear properties in the absence of the thickener [23].
In the classical grease lubrication model, it is usually assumed that “bled” oil released by shear-degradation of bulk grease replenishes the track and forms the lubricating film [24] [25]. However, this model was developed for bearings operating at fairly high speeds and either unidirectional or long amplitude reciprocation. An alternative model suggested by Scarlett [26] was that grease deposits a high-viscosity thickener layer in the track. Evidence of both mechanisms has been demonstrated in the laboratory in ball-on-disc tests [25], [27]. For slow-speed fretting contacts, the entrainment and retention of thickener is likely to be a factor. Under these conditions, the formation of a hydrodynamic film is inhibited both due to the low speed (hc∼ entrainment speed0.7) and the fretting condition. The fretting condition means provision of fluid is restricted as, at stroke reversal, the contact moves back into the depleted exit region. Lubricant replenishment close to the contact, in the absence of excess fluid, is driven primarily by capillary and surface-tension forces [28][29]. Although released oil was present as a thin (estimate ∼µm) film in the scars for high-wear tests, this suggests the base oil alone does not provide sufficient surface protection under these conditions. It is also probable that S/P additives do not work at very low temperatures, as these elements were not detected in the minus-40°C EDS film analysis (Figure 11) but were present at higher temperatures (Figure 10). Reflow-driven replenishment is reduced for high-base oil viscosities and grease [29][30], and this will be exacerbated at low-temperatures. The EDS analysis also showed higher carbon levels for grease A compared to grease D at low temperatures, which correlates with low wear and supports the conclusion that the thickener contributes to load-carrying.
There is an apparent contradiction in this statement as low wear correlated with the lower thickener-content greases. In ball-on-disc unidirectional tests, entrainment of bulk grease into the contact produces much thicker lubricant films than predicted from the base oil viscosity, particularly at low speeds, and has been demonstrated for a range of thickener types [26], [27], [31]. Thus, the thickener will contribute to enhanced load-carrying for low-temperature fretting where the hydrodynamic film is restricted and additive action inhibited.
To explain the low wear/low thickener content contradiction, it is worth considering grease lubrication mechanisms in the fretting regime. At the start of the HFRR test, the ball is loaded onto a 50 µm thick grease film; some grease will be trapped in the contact, but most will be pushed up around the ball. As reciprocation starts, grease in the contact is rapidly broken down; this is often seen as an increase and then rapid decrease in ECR. Bulk grease closest to the ball will be displaced by the reciprocating action and shear-degraded releasing mobile fluid [25] facilitating contact replenishment. The presence of released base-oil in the track has been shown for some of the low-temperature tests. We suggest this is an underlying mechanism, which occurs for all conditions, but it is augmented by bulk-grease replenishment in the low-wear tests.
For small-amplitude motion, it is feasible that displaced grease adhering to the ball is pulled back into the track during reciprocation. Thus, there is a mechanism to continually supply grease to the track, and this might be related to the bulk properties which include rheology (yield stress, visco-elasticity), adherence and tackiness. The apparent contradiction in the low wear/low thickener conclusion can thus be explained by considering the properties that determine grease retention around the contact rather than the absolute thickener-concentration. It is the ability of the grease to maintain replenishment of the contact that is important, and lower base oil viscosity and thickener content would aid this. Thus, yield stress, visco-elasticity, and adherence/tackiness properties could all play a role. “Tackiness” is the property used to describe the ability of a grease to form “strings” when a small sample is retracted [32], and this might contribute to replenishment.
The “adherence” model of replenishment is supported by the image in Figure 7 (lower images) where grease D at minus-40°C has clearly retracted from around the wear scar. Very high wear (ball scar diameter > 500 µm) and friction (average µ∼0.8) were recorded in this test. Grease A at minus-40°C is still retained close to the contact and recorded much lower wear (ball scar diameter ∼250 µm) and friction (average µ∼0.6). The RA-IRS spectra also showed thickener to be present in the lubricant film. As such, it contributes to the low wear but relatively high friction measured for grease A.
A key question is whether these results and proposed lubrication mechanism are applicable to blade pitch bearing operation. The HFRR test conditions of low frequency, restricted stroke amplitude, and high pressure replicate those found in practice. In bearings, there are additional mechanical forces that contribute to replenishment of the track; these include cage effects [33], lateral vibration [34], and transient loading [35][36]. All of these can aid bulk grease movement; however, in the confined fretting motion, it is possible that the adherence mechanism also operates.
The results presented in this article emphasize the complexity of grease lubrication of fretting contacts at low temperatures. Grease lubrication performance is strongly linked to contact replenishment, which depends critically on the test configuration and operation. In fretting tests, we suggest lubricant replenishment is a mixture of released-oil reflow and bulk grease adherence. The efficiency of the reflow mechanism will be determined by the amount of fluid released during shear and the viscosity/polarity of the fluid. Increasing base oil viscosity might increase fully-flooded EHL film thickness, and hence reduce wear, but if it hinders lubricant replenishment of the contact, the inverse occurs. In the current work, increasing base oil viscosity was linked to higher wear and incorporating a low viscosity; polar component in the base oil might facilitate replenishment or low friction behavior.
4 Conclusions
Fretting tests have been carried out to measure the friction and wear performance of four commercial wind turbine greases, for a range of temperatures (25, 7, minus-20, minus-40°C). Similar friction and wear performance was recorded for all the greases at the higher temperatures (25 and 7°C). At the lowest temperature, greases C and D gave markedly poorer results. Overall, for the temperature range tested, the best friction and wear results were for the grease formulated with the lowest viscosity base oil. In low-temperature, fretting contacts base oil replenishment does not provide sufficient surface protection. The presence of thickener in the track also appears to contribute to wear reduction, which suggests bulk grease must be periodically supplied to, or retained in, the track. The “adherence” replenishment mechanism will be linked to the grease rheology and/or tackiness, which will be critically temperature dependent. The contribution of S/P anti-wear additives at low temperatures is unclear and requires further investigation.
Statement of originality
The research work described in this paper is, to the best of the authors’ knowledge and belief, original, except as referenced and acknowledged in the text. The work has not been submitted elsewhere, either in whole or in part, for publication.
Declaration of Competing Interest
The authors declare that they have no known competing financial interests or personal relationships that could have appeared to influence the work reported in this paper.
Data Availability
Data will be made available on request.
References
https://wwindea.org/world-market-for-wind-power-saw-another-record-year-in-2021–973-gigawatt-of-new-capacity-added; accessed 26 September 2022.
A. Dhanola, H.C. Garg; Tribological challenges and advancements in wind turbine bearings: a review. Eng Fail Anal, 118 (2020), Article 104885.
M. Stammler, P. Thomas, A. Reuter, F. Schwack, G. Poll; Effect of load reduction mechanisms on loads and blade bearing movements of wind turbines; Wind Energy, 23 (2020), pp. 274-290.
M. Stammler, A. Reuter, G. Poll; Cycle counting of roller bearing oscillations – case study of wind turbine individual pitching system; Renew Energy Focus, 25 (2018), pp. 40-47.
A.M. Kirk, P.H. Shipway, W. Sun, C.J. Bennett; Debris development in fretting contacts – debris particles and debris beds; Tribology Int, 149 (2020), Article 105592.
P.M. Cann; Thin-film grease lubrication; Proc Inst Mech Eng, Part J: J Eng Tribology, 213 (5) (1999), pp. 405-416.
ASTM D4170: Standard Test Method for Fretting Wear Protection by Lubricating Greases.
ASTM D7594 – Fretting Wear Resistance of Lubricating Greases Under High Hertzian Contact Pressures Using a High-Frequency, Linear-Oscillation (SRV) Test Machine.
https://issuu.com/kim0824/docs/6_nov_dec_2020/52 (accessed 11 October 2022).
F. Schwack, N. Bader, J. Leckner, C. Demaille, G. Pol; A study of grease lubricants under wind turbine pitch bearing conditions; Wear, 454–455 (2020), Article 203335.
T. Maruyama, T. Saitou, A. Yokouchi; Differences in mechanisms for fretting wear reduction between oil and grease lubrication; Tribol Trans, 60 (2016), pp. 497-505.
Z.R. Zhou, L. Vincent; Lubrication in fretting — a review; Wear, 225–229 (1999), pp. 962-967.
ASTM D6079-11; Test method for evaluating lubricity of diesel fuels by the high-frequency reciprocating; Rig (HFRR) (2016).
ISO 12156-1; Diesel fuel – assessment of lubricity using the high-frequency reciprocating; Rig (HFRR) (2016).
K.L. Johnson; Contact Mechanics; Cambridge University Press (1985).
P.M. Cann, M.N. Webster, J.P. Doner, V. Wikstrom, P. Lugt; Grease degradation in R0F bearing tests; Tribology Trans, 50 (2007), pp. 187-197.
P.M.E. Cann, A.A. Lubrecht; Bearing performance limits with grease lubrication: the interaction of bearing design, operating conditions and grease properties; J Phys D: Appl Phys, 40 (18) (2007), pp. 5446-5451.
D.L.A. de Faria, S. Venancio Silva, M.T. de Oliveira; Raman microspectroscopy of some iron oxides and oxyhydroxides; J Raman Spectrosc, 28 (11) (1997), pp. 873-878.
J.L.Hernández Viesca, A. Battez, A.R. González, T. Reddyhoff, A. Torres Pérez, H.A. Spikes; Assessing boundary film formation of lubricant additivised with 1-hexyl-3-methylimidazolium tetrafluoroborate using ECR as qualitative indicator; Wear, 269 (2010), pp. 112-117.
M. Varenberg, G. Halperin, I. Etsion; Different aspects of the role of wear debris in fretting wear; Wear, 252 (2002), pp. 902-910.
Hurley, S. and Cann, P.M. IR Spectroscopic Analysis of Grease Lubricant Films in Rolling Contacts, 25th Leeds-Lyon Symposium on Tribology (1999): 589–600.
B. Li, L. Huang, M. Zhong, Z. Wei, J. Li; Electrical and magnetic properties of FeS2 and CuFeS2 nanoplates; RSC Adv, 5 (2015), pp. 91103-91107.
G. Guangteng, H.A. Spikes; Fractionation of liquid lubricants at solid surfaces; Wear, 200 (1–2) (1996), pp. 336-345.
E.R. Booser, D.F. Wilcock; Minimum oil requirements of ball bearings; Lubr Eng, 9 (140–3) (1953), pp. 156-158.
Merieux, J.-S., Hurley, S., Lubrecht, A.A. and Cann, P.M. Shear Degradation of Grease and Base Oil Availability in Starved EHL Lubrication, Proc. 26th Leeds-Lyon Symposium on Tribology, Tribology Series (2000) 38: 581–588.
Scarlett, N.A., Use of grease in rolling bearings Proc. IMechE., (1967) 182 (3A):167–171.
P.M. Cann; Starved grease lubrication of rolling contacts; Tribology Trans, 42 (4) (1999), pp. 867-873.
L. Huang, D. Guo, S.Z. Wen; Starvation and reflow of point contact lubricated with greases of different chemical formulation; Tribology Lett, 55 (2014), pp. 483-492.
B. Jacod, F. Pubilier, P.M.E. Cann, A.A. Lubrecht; An analysis of track replenishment mechanisms in the starved regime; Tribol Ser, 36 (1999), pp. 483-492.
L. Gershuni, G. Mats, M.G. Larson, P.M. Lugt; Lubricant replenishment in rolling bearing contacts; Tribol Trans, 51 (2008), pp. 643-651.
M. Gao, H. Liang, W. Wang, H. Chen; Oil redistribution and replenishment on stationary bearing inner raceway; Tribol Int, 165 (2022), Article 107315.
S. Achanta, M. Jungk, D. Drees; Characterisation of cohesion, adhesion, and tackiness of lubricating greases using approach–retraction experiments; Tribol Int, 44 (2011), pp. 1127-1133.
B. Damiens, A.A. Lubrecht, P.M. Cann; Influence of cage clearance on bearing lubrication; Tribology Trans, 47 (1) (2004), pp. 2-6.
Y. Nagata, K. Kalogiannis, R. Glovnea; Track replenishment by lateral vibrations in grease-lubricated EHD contacts; Tribol Trans, 55 (2012), pp. 91-98.
P.M. Cann, A.A. Lubrecht; The effect of transient loading on contact replenishment with lubricating greases; Tribology Ser, 43 (2004), pp. 745-750.
X. Zhang, R. Glovnea; An experimental investigation of grease lubricated EHD contact subjected to normal sinusoidally variable loading; Tribol Int, 147 (2020), Article 106272.
Louisiana’s first wind turbine and its components have arrived at Avondale Global Gateway (AGG) after a transatlantic journey from Ireland. Gulf Wind Technology (GWT), headquartered at Avondale Global Gateway in Jefferson Parish, is preparing the onshore turbine for installation at the Port Fourchon Coastal Wetlands Park, with initial deployment and testing slated to begin late this year.
More than 450 local companies are ready to support offshore wind in the Gulf of Mexico. (Courtesy: Gulf Wind Technology)
“This first turbine will demonstrate all the necessary elements for deploying wind-energy projects in the Gulf, marking a crucial step toward realizing the full technical and economic potential for offshore wind,” said James Martin, Gulf Wind Technology CEO. “It’s essentially a prototype to provide us research-oriented results that we can build upon and demonstrate the potential supply chain available in Louisiana, starting with Avondale Global Gateway and finishing at deployment near Port Fourchon.”
“The arrival of this wind turbine underscores Avondale Global Gateway’s commitment to innovation,” said Host Chairman and CEO Adam Anderson. “Avondale is a prime location for companies like Gulf Wind Technology, and we are proud that they call Avondale Global Gateway home. Together, we will continue to increase economic stability and energy development in Jefferson Parish, Southeast Louisiana, and beyond.”
The transport of this turbine tested Louisiana’s pre-built infrastructure that could become part of the offshore wind supply chain. A recent report said more than 450 local companies are ready to support offshore wind in the Gulf of Mexico.
In addition to importing large offshore wind components, Avondale Global Gateway’s modernized enhancements can offer storage, sub-assembly, and on-site manufacturing and fabrication before loading turbine components onto barges for installation in the Gulf. Avondale Global Gateway’s all-encompassing value and proximity to the Gulf’s experienced workforce make it well-positioned to serve as a logistics and supply chain hub for future offshore wind opportunities.
This article reviews the latest developments of substructures for offshore wind turbines focusing on investigations and applications of hybrid foundations. Model tests and numerical analyses were used to simulate the loading of hybrid piles in sand. The results of pile-soil interaction were investigated to confirm the changes in soil stiffness around the hybrid monopile head. The mechanism and factors affecting the change in lateral stiffness of the hybrid foundation were explained by analyzing p–y curves for M+H loading conditions in sand. Based on this research, a new shape of p–y curves for hybrid monopiles was established and a method for determining key parameters was proposed. The effectiveness of new p–y curves was verified by comparing back-calculated results with those from numerical simulations. The conducted tests confirmed that the hybrid monopile displacement is 30 to 50 percent smaller when compared to a standard monopile with similar dimensions. The gained experiences can be useful for designers and researchers to enhance the design of foundations for offshore wind turbines.
Foundation of a windturbine in shallow water, Ijsselmeer, The Netherlands
1 Introduction
Development of wind energy has a major impact on a sustainable, long-term energy balance and on an increasing technological potential. It can be achieved, in short term, through faster development of offshore wind energy. Currently, offshore wind contributes to 45 percent of the total wind capacity installed in Europe. In 2019, 3.6 GW of new capacity was connected to the grid, which is a 1.3-fold increase in capacity comparing to 2018 [1].
Notably, energy obtained from offshore wind shows promise due to higher wind speed and lower disturbance to human lives. Comparing to onshore counterparts, offshore wind has 1.2 to 2 times higher wind speed. In open sea areas, electrical output is expected to be 1.7 times for the same wind turbine, and the energy field tends to be more efficient by going farther from coastlines. Figure 1 shows an evolution of offshore wind turbine sizes. Increasing diameters of rotors transfer increasingly higher loads to turbine foundations.
Figure 1: Evolution of offshore wind turbine sizes.
Additionally, a distance to land for newly installed offshore wind turbines (OWTs) could increase from 30 kilometers to 60 kilometers. Therefore, solid and more stable support structures for new OWTs are needed. There are ambitious plans to increase the share of renewable energy sources by almost 30 percent in the total energy balance all over Europe.
Figure 2: Various ways to shape hybrid foundations for offshore wind turbines.
As a result, up to several hundred new OWTs are planned to be built in Poland between 2027 and 2035 in the southern Baltic Sea. Today, an important challenge for offshore wind energy is to design efficient and reliable offshore wind turbines. The cost of offshore wind support structures (design, construction, and installation) accounts for as much as 35 percent of the investment cost. The installation cost of tower support structures is 60 percent of the total cost for installing a whole wind turbine [2, 3]. Meanwhile, foundation parts have a great role in reductions of the total cost, with the potential of being 6 percent less by introducing innovative monopile techniques. Therefore, reliable and efficient foundations are preferable for the offshore wind industry, and a uniform design of new foundations with the potential of mass production is necessary.
2 Foundation for offshore wind turbines
2.1 Types of foundations
The offshore wind turbine foundations can be divided into main categories depending on the depth of the seabed. Gravity foundations are the right solution for shallow waters (10 to 30 meters). Tripod and jacket foundations are recommended in intermediate waters (30 to 40 meters). Meanwhile, monopile foundations can be installed in waters 40 to 50 meters deep. The concept of floating foundations is best for deep waters (50 to 200 meters). Floating foundations are not often used for commercial OWTs. Currently, the most widely used support structures for OWTs are monopile foundations. They have the largest market share of OWTs in Europe at more than 80 percent.
Figure 3: First applications of hybrid foundations in the world: a) monopile-bucket hybrid foundation [Source: www.offshore-energy.biz], b) hybrid collared monopile [Source: www.rwe.com].
It is anticipated that by 2030, the standard location of wind farms in the sea will be at a depth of 60 meters and 60 kilometers from the mainland. It should be noted that while most operating turbines are supported by monopile foundations, future deployment farther offshore in deeper waters may require more stable structures. Gravity foundations now account for only 5 percent of the market share in Europe. The gravity base provides resistance by its self-weight, and it is fabricated by reinforced concrete with ballasts. Although these materials and construction are less expensive than monopile foundations, the installation cost is a significant concern.
Suction buckets are another popular solution of shallow foundations. They are inserted into the seabed using self-weight and suction, which significantly accelerates the installation process, saving time and cost. However, the application of gravity-base foundations is significantly limited by the water depth and soil conditions. Tripod foundations are structures with a wider base and anchor piles driven to the seabed to hold the foundation firm. Jacket foundations are designed with a lattice truss supported with three or four tubular steel legs. These foundations have high resistance of dynamic responses.
They are used in relatively deeper water; however, their high construction and installation cost is the main limitation for wider practical applications. Monopile foundations are made of large-diameter pipe piles. Limited lateral stiffness and installation cost are disadvantages of this type of foundation. All the presented foundations for offshore wind turbines have their advantages and limitations. Taking this into account, new innovative types of foundations with better technical parameters and wider adaptability have been investigated [4-5].
Figure 4: Load scheme and geometry of hybrid foundation for OWTs.
2.2 Hybrid foundation concept
The novel substructure, “hybrid monopole,” is a new type of foundation proposed to reduce the length of a pile, increase its lateral stiffness, and ease construction in marine conditions [6–9]. Figure 2 presents various ways to shape hybrid foundations.
The main objective of this concept is a reduction in the cost of obtaining wind energy; as turbine sizes get larger, standard monopiles become uneconomical, and thus, there is a need for an alternative new solution. Hybrid monopiles are reliable up to a depth of 45 meters. The support structure consists of a vertical pile and a horizontal plate, which provides extra stiffness against lateral load resistance. The horizontal bearing plate is a circular rigid collar or an element showing the appearance of a bucket. The main benefits of this solution are a shorter pile length and greater lateral stiffness of the foundation.
The horizontal plate provides an additional restoring moment in the pile shaft, and friction under the plate can reduce the lateral movement of the pile foundation. The vertical pressure of the plate acting on the soil in front of the pile provides extra lateral pile resistance. Additionally, the relative scour depth around the pile is reduced, since the plate enlarges the contact area between the foundation and the seabed soil. It helps to minimize the scour failure [10–14]. After the wind turbine’s service life, the plate and pile can easily be removed and decommissioned. There are no codes of practice for design methodology of hybrid monopile foundations. In 2020, the first novel hybrid foundation had been successfully used for supporting an offshore wind turbine at the Putian Pinghai Phase II site on the southeast coast in the Fujian province in China, as shown in Figure 3a.
The hybrid foundation consisted of a large diameter monopile and a wide shallow bucket. For installation, the monopile was first embedded into the seabed and become a guideline for locating a bucket. The bucket was then installed through the center of the monopile by pumping out the water inside the bucket. High-strength grouting materials were then filled in the gap between the monopile and the bucket to connect these two components. The bucket had a diameter of 14 meters and a height of 6.4 meters. The pile was cylindrical and open-ended. Its geometric dimensions were: diameter 6 meters, wall thickness 0.05 meters, and embedding depth 60 meters. Although the hybrid monopile-bucket foundation has been used in practice, the existing studies on the lateral bearing capacity of these solutions are extremely limited [15, 16].
Figure 5: Model tests of piles under lateral load M/H = 4: a), b) test stand and hybrid pile model; c), d) deformation zones near tested pile models; e), f) assumptions of the calculation method for hybrid pile.
Wang [17] performed a series of numerical analyses on hybrid monopile-bucket foundations to investigate its lateral static and dynamic responses. The results of the analysis showed the addition of the bucket to the pile foundation effectively restrained the rotation and lateral displacement. In 2022, a steel hybrid monopile technology premiered at the Kaskasi offshore wind farm on the North Sea in Germany, as shown in Figure 3b. For the first time ever in the renewable energy industry, special collars were installed around the monopiles at the seabed level around three wind turbines [18]. Each wind turbine has a capacity of up to 9 MW. The innovative foundation collars were successfully embedded into the seabed, each 7 meters high, weighting 170 tons. The installation was carried out up to a depth of 25 meters. The space between the collar and the monopile foundation was filled with grout material, firmly connecting these construction elements. The new technology provided additional support for the pile lateral loading, increased the bearing capacity, and improved the structural integrity of this OWT foundation.
3 Research on hybrid foundations
3.1 Scope and purpose
The lateral load bearing capacity of hybrid pile foundations depends mainly on soil stiffness. The requirements for designing OWTs foundations are now more SLS- than ULS-oriented [19- 20]. The serviceability of OWTs may get lost due to excessive tilting or horizontal displacement of the tower at the mudline level. Considering strict operational requirements of turbines,a maximum rotating angle of 0.5 degrees and horizontal displacement of the pile head are often limited. Based on the author’s studies and analyses, a practical method of calculating hybrid monopiles for OWTs was developed. The proposed solution was improved with regards to the pile-soil interaction in the initial phase of increasing displacements. By using the proposed method, calculations for the hybrid monopile allows the impact of a horizontal plate on the monopile to be initially assessed, which generally improves the pile lateral stiffness by 30 to 50 percent. The geometry and the lateral load scheme for hybrid foundations are shown in Figure 4.
Table 1: Soil characteristics in model tests.
3.2 Methodology
In this study, the self-developed monotonic load test loading system was adopted to carry out a series of model tests for a monopile foundation. The theoretical calculation results of lateral load were adopted for the pile-soil interaction analyses and were analyzed in accordance with the model test results. The laboratory model tests were treated as a qualitative assessment of the research problem. The small-scale experiments were conducted at the Rzeszow University of Technology, Poland. Fine silica sand was used for testing. A general view of the test stand and the hybrid pile model are shown in Figure 5a,b.
The box was filled with sand using the curtain method, with the container moving at a constant height to provide uniform compaction conditions. The layers of the sand placed in the box were of different colors for better observation of both models of foundations and soil through the transparent box wall. The soil parameters are summarized in Table 1.
The aim of these tests was to qualitatively assess the behavior of the soil under the plate in front of the hybrid pile laterally loaded. The lateral load scheme was adopted for M/H ratio = 4. The foundation model was subjected to a horizontal force on a reduced scale 1/N2 = 100. The model tests confirmed that standard piles interact with the soil in a different way than hybrid piles. Significant differences in shapes of the active soil zones in the vicinity of two pile models are shown in Figure 5c (standard pile) and Figure 5d (hybrid pile). It was observed that sand movement starts at the beginning of the pile load and continues throughout the load duration. A resistance zone appeared under the plate, which divided the soil area into a convective zone and a wedge zone (ABC). The wedge area is shown in Figure 5e.
Figure 6: Results of the 3D FE analysis: a) numerical model; b) soil wedge range as stress difference for hybrid and standard piles; c), d) impact zones around two types of piles for the same pile head displacement of 20 mm.
When the standard pile was loaded, the entire active area in front of the pile was significantly smaller compared to the hybrid pile with the same displacement. The pivot point zo of the hybrid pile model was found to be in sand approximately 1/3 above that of the standard pile model. Based on the model tests, it was determined the active soil area formed under the bearing plate in front of the hybrid pile is always wedge-shaped. It was assumed on the basis of other studies [21, 22]. The depth of the wedge increases with the hybrid pile loading, as shown in Figure 5f. Based on the previous analyses of hybrid piles, it was found the change of depth of the wedge under the bearing plate occurs only in a limited range of the pile displacement [5, 6]. In this case, the lateral pressure of the pile on the soil is due to the stress wedge. It was assumed that the reaction force R in front of the hybrid pile resulting from stresses in the wedge impact zone depends on the vertical plate pressure and changes with the pile rotation.
3.3 Results
The numerical FE analysis was used to quantify the soil behavior around the hybrid pile with a flexible shaft and to develop a method for calculating its wide displacement range. Numerical modeling was carried out using the FE method and Plaxis 3D software, as shown in Figure 6. The non-linear elastic-plastic Mohr-Coulomb type soil model was used.
Table 2: P–y functions for a hybrid pile in sand.
The analysis was carried out using homogeneous sand with parameters: angle of internal friction 30 degrees, Young’s modulus 80 MPa increased 1.5z with depth (z), cohesion 1 kPa, Poisson’s ratio 0.3, and dilatation angle 5 degrees. The numerical model was validated for the behavior of the standard pile and the experimental data of the field test for a 10 meter-long, 1.2 meter-diameter pile. The presented results of 3D FE calculations are applicable for loads causing the same pile head displacement of 20 mm. In this case, the horizontal force was 400 kN and 1,000 kN for the standard and hybrid piles, respectively. Figure 7 shows the stress-displacement curves of the hybrid and standard piles calculated for different depths.
It can be seen that in the displacement range (y) 10–50 mm, the stress (p) increases significantly for the hybrid pile compared to the standard pile. At depths greater than 2 meters, the curves in both diagrams have a similar shape. The changes of stresses in the soil zone in front of the pile are observable only up to the depth of 2 meters, and their values depend on the displacement of the pile at a given depth. The FE analysis shows the type of p–y curves for depths up to 2 meters is crucial for calculating the hybrid pile. The forms of p–y functions for different depths in the active zone of the soil in front of the hybrid pile can be written by fifth-degree polynomials, as shown in Table 2.
Figure 7: P–y curves for the 10 m-long pile based on FE analysis: red line – hybrid, blue line – standard.
4 Discussion
In engineering practices, special attention should be paid to the actual lateral response of bending hybrid piles in the soil in the zone below the seabed level. For this purpose, a modification of the p–y curves with 2 meters of depth below the plate has been proposed. The new shape of p–y curves is particularly important in the initial range of pile displacements up to 50 mm. Figure 8 shows control calculations for the analyzed standard and hybrid piles.
The assumptions on the shape of new p–y curves for hybrid piles were based on two main considerations: First, numerical analyses show that, for the standard piles, soil failure occurs at a displacement of 10 mm; in the p–y diagram, this corresponds to the value of D/100. Thus, this value can be taken as the upper limit of the range in which the pile is still elastic in the soil. Second, the shape of p–y curves for the hybrid pile is different from that of the standard pile. The main difference is that, above the displacement D/100, the hybrid pile is still stable, unlike the standard pile whose behavior in the ground is clearly non-linear.
Figure 8: P–y curves for piles D = 1.2 m, L = 10 m; a) standard, b) hybrid.
The modification consists in providing new p-stress values for the curves in the range of normalized displacements y in the range from 10 mm (D/100) to 50 mm (3D/80). The p–y curves for the standard and hybrid piles were used in control calculations. According to Figure 9 and Figure 10, the displacement and bending moments of the hybrid pile are smaller than the standard pile.
In both cases, pile displacements and soil response distributions were in a satisfactory agreement with previous FE simulations. Calculations with the modified p–y curves for the hybrid pile allowed for accurate determination of the plate effect Ep. The Ep value described as a ratio of the standard pile displacement to the hybrid pile displacement under the same lateral load was assumed as a measure of the plate effect. The Ep values obtained are consistent with the results of other tests in sand [4, 5]. With an increase in the lateral load of the hybrid pile, the distribution of the maximum bending moment and soil reaction occurs at the same depth. A comparison of the parametric results in Figure 11 shows the lateral stiffness of the analyzed hybrid pile increases with deflection.
When the pile is loaded with a force of 400 kN and a bending moment of 1,600 kN•m, in control calculations, the value of the plate influence coefficient Ep is 1.45. The use of the proposed modified p–y curves affects the horizontal stress distribution in the soil in front of the hybrid pile. As a result of the calculations, lower deflections of the hybrid pile were obtained, which is consistent with the results of the field tests. No significant changes were found in the bending moment distribution in the pile.
Figure 9: Calculation results of the standard pile: a) displacement; b) bending moment.
Calculations also showed a smaller depth of the pivot point of the hybrid pile in the soil. Analyzing the changes in the lateral pressure of the pile on the ground, it can be seen that at a rotation of 0.002 rd, there is no failure of the soil under the plate. This proves the beneficial effect of the plate on the stability of the hybrid pile. The hybrid pile can carry a higher lateral load at the same displacement as the standard pile. The proposed method for calculating hybrid piles is a simplified solution. Nevertheless, the calculated displacement values are generally consistent with the results of accurate numerical calculations. Increasing the accuracy of the proposed method is still possible, but it requires further correction of p–y curves by taking into account, to a greater extent, the contribution of lateral zones around the piles.
Figure 10: Calculation results of the hybrid pile: a) displacement; b) bending moment.
A similar problem was analyzed in the proposal for calculating hybrid piles with rigid and flexible shafts. Details are available in previous studies [7, 23].
5 Conclusions
Currently, solid and more stable support structures are needed for new OWTs. An innovative “hybrid” foundation is a new type of support proposed to reduce the length of a standard monopile, increase its lateral stiffness, and ease construction in offshore conditions. The hybrid monopiles have been analyzed by means of model tests and numerical calculations. The following conclusions are drawn from this study:
1: The stiffness of the hybrid monopile increases with its deflection under lateral loading.
2: The analysis showed that the pile-soil interaction for depths up to 2 meters is crucial for the hybrid monopile.
3.: Modified forms of p–y functions for the hybrid monopile can be proposed for the range of displacements up to 50 mm.
4: The conducted studies confirmed that displacements of hybrid monopiles are 30 to 50 percent smaller compared to displacement of standard monopiles with similar dimensions.
Figure 11: Comparison of lateral stiffness calculation results of hybrid monopiles in sand.
References
WindEurope, “Offshore Wind in Europe, key trends and statistics,” 2021. [Online]. Available: https://wind europe.org/. [Accessed: 15 Feb. 2022].
T. Asim, S.Z. Islam, A. Hemmati, and M.S. Khalid, , “A review of recent advancements in offshore wind turbine technology,” Energies, vol. 15, no. 2, 2022, doi: 10.3390/en15020579.
M. Aleem, S. Bhattacharya, L. Cui, S. Amani, A.R. Salem, and S. Jalbi, “Load utilisation ratio of monopiles supporting offshore wind turbines: Formulation and examples from European wind farms,” Ocean Engineering, vol. 248, 2022, doi: 10.1016/j.oceaneng.2022.110798.
X. Wang, X. Zeng, X. Yang, and J. Li, “Feasibility study of offshore wind turbines with hybrid monopile foundation based on centrifuge modeling,” Applied Energy, vol. 209, pp. 127–139, 2018, doi: 10.1016/j.apenergy. 2017.10.107.
K. Trojnar, ”Lateral stiffness of hybrid foundations: field investigations and 3D FEM analysis,” Geotechnique, vol. 63, no. 5, pp. 355–367, 2013, doi: 10.1680/geot.9.P.0778.
K. Trojnar, “Multi scale studies of the new hybrid foundations for offshore wind turbines,” Ocean Engineering, vol. 192, 2019, doi: 10.1016/j.oceaneng.2019.106506.
K. Trojnar, “Simplified design of new hybrid monopile foundations for offshore wind turbines,” Ocean Engineering, vol. 219, 2021, doi: 10.1016/j.oceaneng.2020.108046.
K.B.M. Lehane, B. Pedram, J.A. Doherty, and W. Powrie, “Improved performance of monopiles when combined with footings for tower foundations in Sand”, Journal of Geotechechnic and Geoenvironment Engineering, vol. 140, no. 7, 2014, doi: 10.1061/(ASCE)GT.1943-5606.0001109.
D. Chen, P. Gao, S. Huang, C. Li, and X. Yu, “Static and dynamic loading behavior of a hybrid foundation for offshore wind turbines,” Marine Structures, vol. 71, 2020, doi: 10.1016/j.marstruc.2020.102727.
F. Liang, C. Wang, and X. Yu, “Widths, types, and configurations: influences on scour behaviors of bridge foundations in non-cohesive soils,” Marine Georesources & Geotechnology, vol. 37, no. 5, 2019, doi: 1080/ 1064119X.2018.1460644.
W.G. Qi, F. Gao, M. F. Randolph, and B. M. Lehane, “Scour effects on p–y curves for shallowly embedded piles in sand,” Geotechnique, vol. 66, no. 8, pp. 648–660, 2016, doi: 10.1680/jgeot.15.P.157.
Z. Wang, R. Hu, H. Leng, H. Liu, Y. Bai, and W. Lu, “ Deformation analysis of large diameter monopiles of offshore wind turbines under scour,” Applied Sciences, vol. 10, no. 21, 2020, doi: 10.3390/app10217579.
H. Ma and C. Chen, “Scour protection assessment of monopile foundation design for offshore wind turbines,” Ocean Engineering, vol. 231, 2021, doi: 10.1016/j.oceaneng.2021.109083.
S. Bajkowski, M. Kiraga, and J. Urbański, “ Engineering forecasting of the local scour around the circular bridge pier on the basis of experiments,” Archives of Civil Engineering, vol. 67, no. 3, pp. 469–488, 2021, doi: 10.24425/ace.2021.138066.
A. Buljan, “First monopile-caisson hybrid foundation installed at Chinese offshore wind farm,” Offshore Energy Project News, 2020. [Online]. Available: https://www.offshore-energy.biz/first-monopile-caisson-hybrid- foundation-installed-at-chinese-offshore-wind-farm/. [Accessed: 15 Jul. 2020].
L.X. Xiong, H.J. Chen, Z.Y. Xu, and C.H. Yang, “Numerical simulations of horizontal bearing performances of step-tapered piles,” Archives of Civil Engineering, vol. 67, no. 3, pp. 43–60, 2021, doi: 10.24425/ace.2021. 138042.
J. Wang, G. Sun, G. Chen, and X. Yang, “Finite element analyses of improved lateral performance of monopile when combined with bucket foundation for offshore wind turbines,” Applied Ocean Research, vol. 111, 2021, doi: 10.1016/j.apor.2021.102647.
S. Knauber, “World’s first: Innovative steel collars installed at RWE’s Kaskasi wind farm in German North Sea,” RWE Renewables. [Online]. Available: https://www.rwe.com/en/press/rwe-renewables/2022-06-08-innovative- steel-collars-installed-at-rwes-kaskasi-wind-farm. [Accessed: 18. Jan. 2023].
API, “Recommended Practice 2A. Planning, Designing, and Constructing Fixed Offshore,” 2014. [Online]. Available: https://api.org/pubs/. [Accessed: 18. Jan. 2023].
B. Yuan, M. Sun, Y. Wang, and L. Zhai, “Full 3D Displacement measuring system for 3D displacement field of soil around a laterally loaded pile in transparent soil,” International Journal of Geomechanics, vol. 19, no. 5, 2019, doi: 10.1061/(ASCE)GM.1943-5622.0001409.
L. Li, X. Liu, H. Liu, W. Wu, B. M. Lehane, G. Jiang, and M. Xu, “Experimental and numerical study on the static lateral performance of monopile and hybrid pile foundation,” Ocean Engineering, vol. 255, 2022, doi: 10.1016/j.oceaneng.2022.111461.
K. Trojnar, “Experimental and numerical investigation of lateral loaded flexible hybrid piles in sand,” International Journal of Civil Engineering, vol. 21, no. 1, pp. 1–18, 2023, doi: 10.1007/s40999-022-00736-x
Amsoil Inc., a leader in synthetic lubricant technology, has acquired Columbus, Ohio-based Aerospace Lubricants. Founded in 1973, Aerospace designs and manufactures a specialized array of greases for industrial manufacturing operations and private-label customers in automotive, industrial, military, aerospace and consumer markets. Aerospace Lubricants will operate as an independent subsidiary of Amsoil.
“By acquiring Aerospace Lubricants, we are acquiring a strong new partner for grease formulation and production,” said Amsoil Chairman & CEO Alan Amatuzio. “We are going to invest in Aerospace to enhance the capabilities and capacity of the operation. Our operational excellence combined with their grease expertise will result in significant value added for Aerospace and AMSOIL customers.”
Aerospace will continue serving existing and new customers in its target markets and will now be led by Dave Meyer, former Amsoil Sr. VP, Industrial, who will assume the role of Aerospace Lubricants President.
“Aerospace Lubricants will continue operating as Aerospace Lubricants, an independent subsidiary of Amsoil,” Meyer said. “It will not be rebranded as Amsoil, but it will have the backing of the Amsoil team and resources to drive operational improvements that will deliver significant benefits for Aerospace customers. I am looking forward to working with the Aerospace team to grow the business together.”
Amsoil and Aerospace are both family-owned companies with long histories engineering innovative lubricant solutions. Both are known for advanced formulations and a focus on developing quality, specialized products that deliver the performance customers have been promised and the satisfaction they deserve.
“This was a strategic choice on my part,” said outgoing Aerospace Lubricants owner Steve Gates. “Aerospace is a family-owned company with a great team and strong values. It was important to me to find new ownership that would uphold our core values and amplify the opportunities for the Aerospace team and our customers. Amsoil is the perfect choice. This is an exciting change for everyone involved.”
The Oceantic Network, a leading organization working to advance offshore wind and other ocean renewable industries and supply chains, applauds the Delaware Legislature’s commitment to the development of offshore wind and advancing to the decarbonization goals outlined in the Climate Change Solutions Act of 2023.
Senate Bill 265, referred to as the Delaware Energy Solutions Act of 2024, allows up to 1,200 MW of offshore wind-energy procurement, encourages regional cooperation, and, most importantly, includes provisions that will allow for streamlined development of onshore transmission, thus building a stronger regional market from which the state will benefit. The bill now heads to Gov. John Carney, who previously pledged his support.“Passage of the Delaware Energy Solutions Act represents a pivotal moment for Delaware and an important milestone for the offshore wind industry up and down the East Coast,” said Sam Salustro, vice president of strategic communications at Oceantic Network.
When signed into law, Delaware will join fellow East Coast states as an active buyer of offshore wind-power generation. Passage will also allow the state to capture more of the economic benefits sparked by offshore wind development, including well-paying jobs associated with the area’s developing offshore wind supply chain. Delaware residents may have already witnessed massive wind-turbine components sailing up the Delaware River, aided by local pilots, for final fabrication in New Jersey. These maritime jobs will only grow as a major wind port finishes construction in the Delaware Bay, and Maryland and New Jersey’s first offshore wind projects begin construction. The University of Delaware has been a leader in offshore wind policy and today provides the industry with well-trained graduates. And Crystal Steel Fabricators, headquartered in Delmar, emerged as an early and important steel provider with components coming out of its Eastern Shore facility in Federalsburg, Maryland.
“Delaware has always been an offshore wind pioneer and critical thought leader in the industry, “ Salustro said. “The state now takes its rightful place as an active offshore wind state ready to play an important role supporting development of the regional supply chain. Oceantic Network applauds Senator Hansen and Delaware’s stakeholders and advocates for advancing this legislation and encourages Governor Carney to fulfill his commitment to sign the bill. The Network looks forward to continuing to support the state in building a strong supply chain.”
NeXtWind, provider of climate infrastructure for renewable energy in Germany, achieved milestones ahead of schedule in the first half of 2024. Within 12 months, NeXtWind quadrupled the optimized repowering generation capacity in its portfolio to 1 GW. This leap was made possible by the acquisition of additional wind farms, the optimization of existing wind-farm sites and the technical advancement of the latest generation of turbines.
NeXtWind focuses on acquiring existing wind-energy sites with outdated turbines, replacing the turbines with more efficient ones. (Courtesy: NeXtWind)
NeXtWind focuses on acquiring existing wind-energy sites with outdated turbines, replacing the turbines with more efficient ones, and expanding the sites through cooperation with the local communities to install additional turbines on the site. The company operates 158 wind turbines in 24 wind farms. After optimization, those wind farms will produce four times as much energy, enough to supply more than 600,000 homes with green electricity every year. The current regional focus is on the north and east of Germany, but the aim is for future expansion beyond national borders. The company works closely with local communities and partners to ensure a sustainable and reliable energy supply, through which it is making a significant contribution to the energy transition.
“We originally wanted to reach the magic threshold of 1,000 MW of optimization potential by 2026,” said Prof. Dr. Werner Suess, Co-CEO and Co-Founder of NeXtWind. “The fact that we have now achieved this two years ahead of our goal is very encouraging. NeXtWind is not only a leader in the field of repowering, but also an overall leader in the optimization of the distributed, climate-friendly energy infrastructure in Germany. We will continue to pursue our strategic direction, consolidate our market position, and grow rapidly.”
“NeXtWind is building the next-generation energy infrastructure to drive the transition to a decarbonized, decentralized, and digitized energy system in the face of unprecedented demand for clean, sustainable energy from energy-hungry applications such as AI, blockchain, and the electrification of the mobility and heating sectors,” said Lars Meyer, co-CEO and co-founder of NeXtWind. “We are just getting started — our goal is to shape the energy transition in the coming decades,” said Ewald Woste, Executive Chairman and Co-Founder of NeXtWind. “We aim to triple our generation potential to 3,000 megawatts by 2028.” The NeXtWind team has doubled to more than 50 employees, and the Berlin office has been expanded. Further acquisitions are on the agenda for the coming months.
In a Request for Proposals (RFP), Dominion Energy is seeking power purchase agreements from renewable and other carbon-free energy sources in a region that includes 12 Mid-Atlantic states and the District of Columbia.
Dominion Energy is seeking power purchase agreements from sources in a region that includes 12 Mid-Atlantic states and the District of Columbia. (Courtesy: Dominion Energy)
DEV will only consider proposals for facilities within PJM territory (Pennsylvania-New Jersey-Maryland Interconnection, LLC), but not including facilities in the state of Virginia.
All electrical output from the facilities will be delivered to the PJM Dominion Transmission Zone. Virginia Electric and Power Company is a wholly owned subsidiary of Dominion Energy and is a regulated public utility that generates, transmits, and distributes electricity for sale in Virginia and portions of northeast North Carolina.
Facilities that achieved a commercial operations date (COD) after October 1, 2021, and facilities under construction that achieve COD prior to the end of calendar year 2035 are eligible.
All participating bidders must register by submitting an Intent to Bid Form and an executed Confidentiality Agreement no later than August 30. The Intent to Bid Form, CA and other additional information on this RFP can also be found on the company’s website. The completed form and signed CA should be emailed to DEVCarbonFreeRFP@dominionenergy.com.
Technip Energies and SBM Offshore announce the formal implementation of Ekwil, a 50/50 floating offshore wind joint venture.
Ekwil is a pure player delivery partner offering a diversified range of series production floating offshore wind solutions to meet the growing and demanding needs of energy customers around the world. Ekwil brings together expertise and experience of two energy transition leaders to collectively power progress with semi-submersible INO by Technip Energies and Tension Leg Platform Float4Wind by SBM Offshore. This approach covers a large spectrum of the FOW market, aiming to bring these technologies to commercial deployment.
SBM and Technip Energies are teaming up for an offshore wind joint venture. (Courtesy: SBM Offshore)
Headquartered in France, Ekwil relies on a core team of 40 specialists, bringing together knowledge and innovation capacities and will be backed by the resources of SBM Offshore and Technip Energies for project execution.
“By bringing together two world leading players, Ekwil will accelerate the deployment of industrial solutions for the nascent Floating Offshore Wind market,” said Arnaud Pieton, Technip Energies CEO. “This joint-venture with SBM Offshore illustrates the commitment of Technip Energies to provide a diversified and expanding range of low-carbon solutions to support the global net-zero trajectory.”
“It’s just a question of time for market potential in Floating Offshore Wind power to materialize,” said Øivind Tangen, CEO of SBM Offshore. “This collaboration with Technip Energies ensures the availability of optimal solutions with certainty and reliability in delivery. Ekwil leads both partners toward success, pioneering new standards in renewable energy and driving progress towards a net-zero future.”
With 25 years of experience in the offshore industry, Séverine Baudic, former managing director of New Energies & Services at SBM Offshore, is the CEO of Ekwil.
“Today’s launch of Ekwil marks a significant step to power progress in the floating offshore wind market, combining industry-leading expertise and solutions,” Baudic said. “I am proud to have the trust and commitment of SBM Offshore and Technip Energies and look forward to leading our talented teams towards a greener future for all.”
