STRUCTURE magazine - January 2018

January 2018 Concrete

Inside: Mountain “S” Home, Utah

A Joint Publication of NCSEA | CASE | SEI

STRUCTURE ®

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LeMessurier Calls on Tekla Structural Designer for Complex Projects Interoperability and Time Saving Tools

Tekla Structural Designer was developed specifically to maximize collaboration with other project parties, including technicians, fabricators and architects. Its unique functionality enables engineers to integrate the physical design model seamlessly with Tekla Structures or Autodesk Revit, and to round-trip without compromising vital design data.

“We’re able to import geometry from Revit, design in Tekla Structural Designer and export that information for import back into Revit. If an architect makes geometry updates or changes a slab edge, we’ll send those changes back into Tekla Structural Designer, rerun the analysis and design, and push updated design information back into Revit.”

Tekla Structural Design at Work: The Hub on Causeway

For over 55 years, LeMessurier has provided structural engineering services to architects, owners, contractors, developers and artists. Led by the example of legendary structural engineer and founder William LeMessurier, LeMessurier provides the expertise for some of the world’s most elegant and sophisticated designs while remaining true to the enduring laws of science and engineering.

Known for pushing the envelope of the latest technologies and even inventing new ones, LeMessurier engineers solutions responsive to their clients’ visions and reflective of their experience.

An early adopter of technology to improve their designs and workflow, LeMessurier put its own talent to work in the eighties to develop a software solution that did not exist commercially at the time. Their early application adopted the concept of Building Information Modeling (BIM) long before it emerged decades later.

While LeMessurier’s proprietary tool had evolved over three decades into a powerhouse of capability, the decision to evaluate commercial structural design tools was predicated on the looming effort required to modernize its software to leverage emerging platforms, support normalized data structure integration and keep up with code changes.

After a lengthy and thorough comparison of commercial tools that would “fill the shoes” and stack up to the company’s proprietary tool, LeMessurier chose Tekla Structural Designer for its rich capabilities that addressed all of their workflow needs. According to Derek Barnes, Associate at LeMessurier, ” Tekla Structural Designer offered the most features and the best integration of all the products we tested. They also offered us the ability to work closely with their development group to ensure we were getting the most out of the software.”

One Model for Structural Analysis & Design

From Schematic Design through Construction Documents, Tekla Structural Designer allows LeMessurier engineers to work from one single model for structural analysis and design, improving efficiency, workflow, and ultimately saving time. “Our engineers are working more efficiently because they don’t need to switch between multiple software packages for concrete and steel design. Tekla Structural Designer offers better integration of multiple materials than we have seen in any other product,” said Barnes.

LeMessurier engineers use Tekla Structural Designer to create physical, information-rich models that contain the intelligence they need to automate the design of significant portions of their structures and efficiently manage project changes.

TRANSFORMING THE WAY THE WORLD WORKS

“Tekla Structural Designer has streamlined our design process,” said Craig Blanchet, P.E., Vice President of LeMessurier. “Because some of our engineers are no longer doubling as software developers, it allows us to focus their talents on leveraging the features of the software to our advantage. Had we not chosen to adopt Tekla Structural Designer, we would have needed to bring on new staff to update and maintain our in-house software. So Tekla Structural Designer is not just saving us time on projects, it is also saving us overhead.

Efficient, Accurate Loading and Analysis

Tekla Structural Designer automatically generates an underlying and highly sophisticated analytical model from the physical model, allowing LeMessurier engineers to focus more on design than on analytical model management. Regardless of a model’s size or complexity, Tekla Structural Designer’s analytical engine accurately computes forces and displacements for use in design and the assessment of building performance.

“Tekla Structural Designer offers better integration of multiple materials than we have seen in any other product.”

Positioning a large scale mixed-use development next to an active arena, a below grade parking garage, and an interstate highway, and bridging it over two active subway tunnels makes planning, phasing and engineering paramount. Currently under construction, The Hub on Causeway Project will be the final piece in the puzzle that is the site of the original Boston Garden.

Despite being new to the software, LeMessurier decided to use Tekla Structural Designer for significant portions of the project. “Relying on a new program for such a big project was obviously a risk for us, but with the potential for time savings and other efficiencies, we jumped right in with Tekla Structural Designer. It forced us to get familiar the software very quickly.”

“Tekla Structural Designer allowed us to design the bulk of Phase 1 in a single model,” said Barnes. The project incorporates both concrete flat slabs and composite concrete and steel floor framing. “Tekla Structural Designer has the ability to calculate effective widths based on the physical model which is a big time saver,” said Barnes. “On this project, the integration with Revit, along with the composite steel design features enabled us to work more efficiently. Adding the ability to do concrete design in the same model was a bonus because we had both construction types in the same building.”

“Tekla Structural Designer helped this project run more efficiently, and in the end it was a positive experience,” said Blanchet.

“Tekla Structural Designer gives us multiple analysis sets to pull from, which gives us lots of control. Most programs don’t have the capability to do FE and grillage chase-down. For the design of beam supported concrete slabs, Tekla Structural Designer allows us to separate the slab stiffness from the beam stiffness, so if we choose to we can design the beams without considering the influence of the slab. In the same model we can use a separate analysis set to review the floor system with the beams and slab engaged,” said Barnes.

Barnes also shared similar benefits with concrete column design. “Tekla Structural Designer does grillage take-downs floor-by-floor, finds the reactions and applies them to the next floor. This allows us to view column results both for the 3-dimensional effects of the structure as a whole and from the more traditional floor-by-floor load take-down point of view.

Doing both has always required significant manual intervention, but Tekla Structural Designer puts it all in one place.” “We reduce the possibility for human error because with Tekla Structural Designer less user input is required,” said Barnes. “Tekla Structural Designer automatically computes many of the design parameters, such as column unbraced lengths. The assumptions made by the software are typically correct, but we can easily review and override them when necessary.”

“Tekla Structural Designer provided the best fit for our workflow compared to other commercially available software.”

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CONTENTS Features

Cover Feature

33 VIRTUAL OUTRIGGERS AND CREATIVE ENGINEERING

By Chris Crilly, P.E., Mark Tamaro, P.E., and Roberto Stark, Ph.D. Design of the tallest building in Mexico hit a snag when wind-tunnel testing identified issues. The solution meant re-thinking and improving the building’s lateral system. An ingenious “virtual outrigger” system reduced the building period by more than 20%.

36 75 ROCKEFELLER PLAZA

30 MOUNTAIN "S" HOME

By John Hinchcliffe, P.E., Joe Mugford, P.E., and Ramon Gilsanz, P.E., S.E. Altering an existing structure's load path is challenging. A composite

By James M. Williams, P.E., C.E., S.E., AIA Every design element of the Mountain “S” Home represents an engineering challenge, from the curving floor plan and 30 individual kite-shaped roofs to perimeter walls made up almost entirely of glass, floating concrete chimneys, suspended floors, and more.

Columns and Departments

box girder, an uncommon element in buildings, was used as the new transfer girder. Three-dimensional finite element models of the existing and modified moment frames were used to verify the modifications.

STRUCTURAL PERFORMANCE

22 An Overview of Fire Protection for Structural Engineers – Part 1

EDITORIAL

7 New Year’s Resolutions

By Corey M. Matsuoka, P.E.

John “Buddy” Showalter, P.E.,

and Richard L. Emberley

David P. Tyree, P.E., and Sandra Hyde, P.E.

STRUCTURAL SYSTEMS

SPOTLIGHT

By Jared S. Hensley, P.E.

9 Post-Tensioned Concrete Construction

51 Connecting Chicago’s “Second Shoreline”

By Kurt J. Naus, P.E., S.E., Matthew F. Hellenthal, P.E., S.E.,

STRUCTURAL DESIGN

By Michael Schwager, P.E., Guido Schwager, P.E., and Marcus Schwager

40 Shrinkage-Compensating Concrete Designs

By Susan Foster-Goodman and Ken Vallens

BUILDING BLOCKS

12 Fire Resistance of Flat Plate Voided Concrete Floor Systems

By David A. Fanella, Ph.D., S.E., P.E., Michael Mota, Ph.D., P.E., SECB, and

44 Post-Installed Adhesive Anchor Systems

46 Structural Engineers and Climate Change

14 Self-Leveling Self-Help

By Beth Lee and Josh Jonsson

By Megan Stringer, S.E. and Mark D. Webster, P.E.

STRUCTURAL PRACTICES

18 Support Restraints and Strength of Post-Tensioned Members – Part 2

BUSINESS PRACTICES

By Jennifer Anderson

STRUCTURAL SUSTAINABILITY

PRACTICAL SOLUTIONS

and Daniel M. Gross, P.E.

58 Looking for a Job? INSIGHTS

By Christopher Gamache, P.E.

Amy Trygestad, P.E.

By Bijan O. Aalami, Ph.D., S.E.

48 2018 IBC and 2018 IEBC Changes Related to Wood Construction

By Frederick W. Mowrer, Ph.D.

26 Wood-Framed Shear Walls

CONSTRUCTION ISSUES

CODES AND STANDARDS

IN EVERY ISSUE 8 Advertiser Index 43 Resource Guide – Anchor Updates 52 NCSEA News 54 SEI Structural Columns 56 CASE in Point

Publication of any article, image, or advertisement in STRUCTURE® magazine does not constitute endorsement by NCSEA, CASE, SEI, the Publisher, or the Editorial Board. Authors, contributors, and advertisers retain sole responsibility for the content of their submissions.

STRUCTURE magazine

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January 2018

Editorial

new trends, new techniques and current industry issues

New Year’s Resolutions

By Corey M. Matsuoka, P.E., Chair CASE Executive Committee

A

ccording to Wikipedia, a New Year’s resolution is a tradition in which a person resolves to change an undesired trait or behavior, to accomplish a personal goal, or otherwise improve their lives. According to Urban Dictionary, a New Year’s resolution is a goal that you propose then forget about the next day. With that in mind, I would like to offer nine New Year’s resolutions to improve your life as a structural engineer. Most of them you already know but, like any good resolution, you just need a gentle reminder.

1) Work Smarter. When I first started working, I complained to my father about the number of hours I had to put in to get the job done. However, I did not get the sympathy I was hoping for.

Instead, he told me, “work smarter, not harder.” Think about three things that you can do more efficiently. Is there something that you are currently doing that could be delegated? Can your e-mail box be decluttered? Are there distractions pulling you off task?

2) Build Your Network. As you have probably heard, it is not what you know; it is whom you know. More and more clients are viewing structural engineering as a commodity (even though I do not agree). What do you do to separate yourself from your competition? For us, it is whom we know. I have had clients tell me they hired us because of the network we have and our ability to navigate complex issues and challenges by picking up the phone.

3) Be Punctual. In high school, I never was able to be on time. I was so bad, my friends would tell me to show up 30 minutes before our actual meeting time, and I would still be late. In college, I became the President of the Hawaii Club at UCLA. I was in charge of our events and needed to be there on time. It would bother me that others would be late and would make me wait for them. It seemed they considered their time was more valuable than mine. From then on, I lived by the mantra “early is on time, on time is late.”

4) Reduce Stress. Regular physical activity reduces the risk of high blood pressure, stroke, and coronary heart disease. On the other hand, sustained stress over time may increase your blood pressure and increases your risk of heart disease. One of my friends worked so hard, he had a heart attack. While recovering in the hospital, a client visited him to go over a proposal. Hopefully, it does not take a heart attack and a client hospital visit for you to realize if you are working too much.

5) Get Out of the Office. There are many benefits to a healthy worklife balance. It enables one to experience both personal satisfaction and professional fulfillment. For me, this means coaching my son’s AYSO soccer team. Twice a week, I am forced out of the office at 4:00 to coach practice, 2 or 3 hours earlier than I would usually leave. Of course, trying to keep ten 8- and 9-year-old boys attentive and organized on a big field with balls might cause me more stress than sitting at my desk (negating the effect of number 4 above), but at least I am getting out of the office.

6) Get Active in a Professional Organization. It is not enough to just join an organization; you need to be engaged. Professional organizations are always looking for volunteers, so it is easy to join a committee or board. In exchange, you will get the opportunity STRUCTURE magazine

to practice leadership skills, learn about the profession, share challenges with like-minded colleagues, enhance your network, and give back to the profession. I can honestly say that, for each organization I have been involved with, I have gotten more out of it than I have put in.

7) Get Educated. It has never been easier to learn new things, as access to knowledge is all around us. STRUCTURE magazine comes to our office, and webinars and the internet are available at our fingertips. I tell interns in orientation, “There is such a thing as a dumb question: If I can Google the answer in 30 seconds, I am sure you can do the same in 15.” If you prefer the old fashioned way, set a goal of finishing a book every other month or set aside 15 minutes a day. Then find ways to implement what you have learned.

8) Improve Relationships. Engineering is a team sport. To excel, one needs to create synergies with colleagues to increase the productivity and effectiveness of the team. A clear deterrent to this is challenging personal relationships. If there is someone in your office that you do not see eye to eye with, take the initiative and meet with him/her to discuss the issues that prevent you from having positive interactions. I know it is tough but, like with taking off a bandage, it is best done quickly and not left to drag out.

9) Welcome/Deliver Bad News. The bad news is not like wine… it does not get better with age. Create a culture that encourages people to communicate bad news and to do so early. Do not live in a firm where fear rules and a shoot-the-messenger philosophy prevails. In these firms, people withhold bad news hoping (like the ostrich with its head in the sand) that the situation will improve.

If bad news is delivered early, firms can take action to mitigate any foreseeable negative impacts.

Now, if you will excuse me, it is time for me to renew that gym membership that I am never going to use again.▪ Corey M. Matsuoka is the Executive Vice-President of SSFM International, Inc. in Honolulu, Hawaii. He is the chair of the CASE Executive Committee. He can be reached at cmatsuoka@ssfm.com.

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January 2018

ADVERTISER INDEX

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Greg Schindler, P.E., S.E. KPFF Consulting Engineers, Seattle, WA Stephen P. Schneider, Ph.D., P.E., S.E. BergerABAM, Vancouver, WA John “Buddy” Showalter, P.E. American Wood Council, Leesburg, VA

January 2018, Volume 25, Number 1 ISSN 1536-4283. Publications Agreement No. 40675118.

STRUCTURE® is owned and published by the National Council of Structural Engineers Associations with a known office of publication of 645 N. Michigan Ave, Suite 540, Chicago, Illinois 60611. Structure is published in cooperation with CASE and SEI monthly. The publication is distributed as a benefit of membership to members of NCSEA, CASE and SEI; the non-member subscription rate is $65/yr domestic; $35/yr student; $90/yr Canada; $60/yr Canadian student; $125/yr foreign; $90/yr foreign student. Application to Mail at Periodical Postage Prices is Pending at Chicago and at additional Mailing offices.

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STRUCTURE® is a registered trademark of National Council of Structural Engineers Associations (NCSEA). Articles may not be reproduced in whole or in part without the written permission of the publisher.

P

ost-tensioning is the most significant development in the concrete construction industry since steel reinforcement was first employed in the mid-1800s.

Post-tensioning (PT) delivers roughly four times the tensile strength compared to conventional reinforcement and significantly reduces (or eliminates) concrete cracking, thus enabling thinner slab construction – reducing the environmental impacts, saving material and labor costs, and shortening construction schedules. Post-tensioning also brings a host of seismic advantages to a structure and enables architects to employ concrete in artful shapes and sizes once thought impossible.

Consequently, post-tensioning in new construction has blossomed in the United States since the 1950s. And, as with any technological advent, the decades following its introduction saw significant improvements in PT techniques and materials. The original button-headed wire with heavy wax paper wrappings has been replaced by higher grade steel strand, excellent anchorages, injection ports, grout caps, polypropylene ducts, improved grouts, and an array of customizable installation techniques to best suit the slab or beam’s design and use.

However, construction concerns arise as structures built in the industry’s formative years, and newer structures for that matter, are remodeled, repurposed, or repaired. The PT remodel industry steadily grows because generations of structures erected using this technology are coming to the ends of their useful lives. These are mostly parking garages and mid-rise buildings erected in the 1960s and 1970s. Construction from the 1970s tends to be particularly “light,” with design engineers saving every foot of rebar and pound of concrete possible, making for underbuilt structures by today’s standards. By the 1990s, PT had been primarily refined into the processes and materials employed today, making remodeling and repairs easier. Also, in previous generations, life-cycle calculations were not given the importance or precision that they are today, so it is little wonder that older buildings require work to extend their usefulness.

Even recently erected structures may require significant work. Perhaps a landlord wishes to add or change utilities in a building or redesign the floor layout. Perhaps a cracked slab has exposed a section of strand, or an exposed anchorage has corroded, or a contractor has cut into a slab and unintentionally severed a PT strand. This article provides an overview of a few common situations and solutions for PT tendon replacement, repair, and remodeling in older and modern commercial structures alike.

Contemporary Concerns Each project brings unique challenges, so begin by consulting an experienced and competent post-tensioning engineer. Simply because one

has successfully repaired or remodeled a building from the same region or date as a previous project does not mean that the prior experience can inform the final details of a new situation.

In fact, it is possible for an older building to show dramatic PT variation within a single building, such as button-headed PT on one level and steel-rod PT on another. Also, without the PT shop-drawings, even in a newer building, the actual placement of the tendons and interactions of the tendons are not known without some investigation informed by experience. So, before any slab or beam is touched, seek professional guidance. What follows are general thoughts based on years of industry experience.

First, slab and beam cracking is a common concern and an indication that attention and repairs are needed. Are the cracks deep? If so, the engineer must ascertain the cause. Often the culprit is not the concrete or post-tensioning work. Most likely, it is the original design. Design issues likely mean

ConstruCtion issues discussion of construction issues and techniques

Post-Tensioned Concrete Construction expansive, expensive repairs rather than a localized and inexpensive solution. To the surprise of many engineers, it is often possible to replace the tendons in existing slabs and beams to increase their strength and close the cracking. One caveat worth mentioning involves epoxy. If the cracks are old, an owner may have filled them with epoxy to improve aesthetics or to protect the slab or beam from water intrusion. However, if the filler epoxy reached the strand, PT replacement is significantly more challenging. A second, clear sign of trouble in an older building is deflection in any region: slab, beams, columns, walls, or ceiling.

Again, the engineer must discover the cause of the deflection and draw up the remodel action plan accordingly. Deflection generally indicates a more serious design issue than cracking.

Parking garage projects account for a great deal of PT work because the technology suits the structures so well. Transportation needs also change over time, requiring garage modifications.

Required live load increases can create challenging situations. It is more likely that a garage will experience live loads near or over capacity than a mid-rise residential building. However, the PT remodel process is usually much easier than in a tenanted building as there is so much working room, and the garage can usually be closed while work is done. Both slabs and beams can be remodeled. Repairs are common too.

Mid-rise buildings with PT slabs rarely experience the live load issues of parking garages.

However, slab repairs required to address broken or damaged tendons, or modifications required

STRUCTURE magazine

9

Tendon Repairs and Modifications

By Michael Schwager, P.E., Guido Schwager, P.E., and Marcus Schwager Michael Schwager is the President of Schwager Davis, Inc. and Schwager Development. He serves on boards and committees for the Post-Tensioning Institute (PTI) and the American Segmental Bridge Institute (ASBI).

He may be reached at mike@schwagerdavis.com.

Guido Schwager is Founder and Chairman of Schwager Davis, Inc. and Schwager Development.

He serves as co-chair of the PTI Committee M-50/ASBI Joint Task Group and is an executive member of ASBI. He may be reached at guido@schwagerdavis.com.

Marcus Schwager is a writer for Schwager Davis, Inc. and Schwager Development. He may be reached at marcusschwager@gmail.com.

for remodeling, are often more challenging to engineer solutions for, primarily because the impact on tenants must be considered.

In general, there are two basic techniques for accomplishing a PT project in a midrise building – clearing a large portion of the building to do all of the work in one phase or performing the work in multiple phases over a more extended period.

The first approach involves shoring three to four stories below the repair floor. This will mean the evacuation of a significant part of the building, interrupting work and rents for the tenants and owner. Also, the shoring itself is costly to erect. The main advantage of this technique is for the PT repair crew since they have free access to the floors and their work can be accomplished quickly. Additionally, if a building has serious design flaws, this technique will probably be the only remodel means available to an owner.

The second approach can be accomplished without traditional shoring. Most structures, even the “lighter” 1970s structures, can usually be remodeled by working on every third strand, one at a time, and proceeding down the slab. This means the crew will usually make at least three passes over the slab, spending significantly more time on the floor than with the shoring method. However, the combined benefit to tenants (they can remain on all other floors) and the crew (they do not need to erect costly shoring) usually makes the added time and labor for the PT crew on the individual floors well worth it, from a cost analysis and tenant relationship standpoint. This method requires evacuation of the remodel floor to clear that floor’s live load. Live load elimination means that a single strand being replaced in any given location will not affect the slab negatively.

Usually, the slab has enough tension reinforcement to absorb the relatively minor, temporary changes in force due to remodel activity. This technique has been employed for applications from simple single-floor, single-strand replacement to multi-floor cutouts for elevator shafts and upgraded escape routes. The amount of actual disruption such remodeling may cause tenants largely depends on the state of the slab. For instance, if the strand is snagging in its ductwork for some reason, small windows to the strand will be jackhammered in the slab to identify and remove the issue, thereby increasing potential tenant disruption.

This process of removing a live load and treating the work one strand at a time probably suggested itself by early repair of failed strands, whether due to crumbling concrete at an anchor head, a slipping head from a

Figure 1. Worker repairs parking garage slab PT strand with couplers.

now-obsolete coil anchor, or an unintentional cut through a strand from a trade worker. In these cases, the tension from one strand was lost, but the slab did not fail because building specifications called for at least that much extra carrying capacity, even in older structures. It is readily inferred, then, that one may approach a repair with a similar mindset, one strand at a time. However, if a strand fails or is unintentionally cut, seek professional guidance.

It cannot be left cut or un-tensioned simply because one sees no cracking, buckling, or other negative signs in the slab. The force changes will have an effect over time, and it is essential that the slab tension capacity is put back into balance to ensure safe and consistent operation for present and future use of the structure. If one discovers deflections of any kind (sagging, buckling, leaning, etc.) in a building, consult an engineer. If the deflection occurs while a crew is working or people are in the building, the building must be evacuated immediately until an engineer investigates the situation and clears the building for reoccupation. This, of course, is true of buildings constructed with or without post-tensioning.

Example Repair: Simple Repair Often the best approaches to forming a sounder general understanding of a subject are to consider specific cases. First, consider a common issue: the repair of a single broken tendon, and second: the creation of a new opening as part of a remodel project.

Typically, commercial buildings are built with 0.5-inch diameter, 7-wire monostrand tendons. Modern materials and techniques make most unbonded tendon repair or replacement a routine event. Older buttonheaded tendons usually feature grease and

STRUCTURE magazine

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January 2018

paper conduit. This conduit is known for inconsistencies that make tendon replacement difficult and, therefore, most button-headed repairs will be splices. If button-headed tendon replacement is required, it may only be possible to run a smaller gauge tendon in its place, and engineers will need to be consulted to ensure that the result remains safe and balanced. In some cases, the new strand may be welded onto the end of the old tendon and pulled through, enabling a 0.5inch diameter monostrand tendon to replace a 0.5-inch diameter button-headed tendon, despite some conduit issues.

Finding the break is usually not a problem, even when hidden in the slab. A scan with a noninvasive ground penetrating radar device provides a wealth of information concerning the location, size, and material of all slabencased elements. Also, if the broken tendon is to be pulled out for removal, the lengths of each segment can be determined to verify the break location.

The first step in construction is to locally shore the work area and to chip away the concrete to expose the broken tendons (Figure 1). This must be done with care to avoid damaging adjacent tendons. In most situations, a single broken tendon will be repaired with a splice and re-tensioned rather than entirely replaced. However, if the tendon shows rust or other material damage, it will need replacement.

Example Repair: Complicated Case If a single-strand repair or replacement represents the basic end of the repair spectrum, a large, multi-floor slab cutout for a new stairway or elevator may serve as the more complicated and challenging end of

Figure 2. PT worker detensioning a tendon to prepare for strand replacement.

the spectrum. A few warnings should be heeded. First, plan on tenant evacuation and building shoring. Retrofits of this nature require open access to ensure safe and timely project completion. Second, review the original plans and scan the slabs to avoid locations of bundled tendons.

Third, even if extensive shoring is temporarily taking load from any columns, avoid destructive construction techniques near them. Besides the obvious need to protect essential building structures, tendon work near columns poses two risks: 1) tendons are often bundled near columns, and bundled tendons are more challenging to work with than single tendons, and 2) tendons are often nearest the top of the slab surface at column locations. Because splice and anchor retrofits need to be fully buried in the slab once the work is complete, much more of the slab must be broken out when repairing any tendon that approaches the surface of the slab. Such tendons must be re-profiled by breaking out and excavating under a tendon, resulting in a redistribution of stress and incurring extensive slab demolition.

As mentioned above, tendons are more challenging to repair when bundled. Therefore, when possible, engineers should avoid bundled locations when selecting a work location,

Figure 3. Slab opening with antiquated button-headed wire system.

especially when working with a multi-floor cutout. Sometimes surrounding tendons will pinch a tendon in the bundle and hinder the work. The tendon may bind when being detensioned. The conduit may collapse when the tendon is removed (again, welding on a new tendon to the old for pull-through replacement may prevent this). Any work done at a bundle location runs a higher risk of damage to other tendons and of increasing labor, material costs, and project-completion time.

Once planning, blocking, de-tensioning, cutting, and demolition are complete, the tendons are replaced or spliced with appropriate anchors, in a manner similar to that for the single tendon project, and the slab is reconstructed. Figure 2 shows the de-tensioning of the tendons at the end of a beam. The process of slab tendon de-tensioning is similar.

A safety item to note is that slab ends must be guarded with perimeter blocking (typically with a heavy wood or steel beam) when the tendons are first cut. The force released by a cut tendon may cause the tendon to break through its grout cap and pose a threat to nearby people or equipment. The contractor must ensure that, before any cutting takes place, safety blocking protects any strand ends that might release force. Once the detensioning and demolition are complete, the

Figure 4. Worker re-tensions a repaired strand.

profile of the tendons can also be adjusted within the work area to address the new structural spans and geometry. Once the tendon work is finished, including the addition of new end anchorages, the concrete floor slab is recast with rebar reinforcement (Figure 3).

Bonding agents help rapid set repair mortars tie into the existing slab. The slab opening may call for a turned down beam to rim the cutout for reinforcement, aesthetic look, or other mounting needs (such as serving as an anchor for railing attachments). Once the concrete cures, tendons are re-tensioned as in new construction (Figure 4), and the cutout is ready for new service (Figure 5).

Conclusion Hopefully, these brief thoughts and examples will serve to clarify basic PT repair and remodeling processes. Like the development of many industries since the 1950s, construction methods and materials have refined dramatically, even within one of the most ancient building materials on earth: concrete. The post-tensioning skill and experience of engineers, crew leaders, and crews means that new and renovated posttensioned structures can grow in scale, number, and elegance.â–Ş

Figure 5. A successful PT slab cutout for a new staircase.

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January 2018

Building Blocks updates and information on structural materials

F

lat plate voided concrete slab systems have been used for many years in Europe and other parts of the world. These systems are becoming increasingly popular in the U.S. This is due to many inherent benefits which include reduced self-weight (resulting in smaller column sizes and foundations as well as smaller seismic forces); larger allowable superimposed loads for given span lengths; economical longer spans; reduced floor-to-floor heights; and accelerated construction schedules.

Contemporary systems, pioneered by BubbleDeck® and Cobiax®, utilize hollow, plastic balls (commonly referred to as void formers) made of high-density, recycled polyethylene (HDPE) that are regularly spaced within the overall thickness of the concrete slab. Void formers are usually spherical or ellipsoidal and are positioned within wire support cages to create modular grids (cage modules), which are locked between the upper and lower reinforcement

Fire Resistance of Flat Plate Voided Concrete Floor Systems

By David A. Fanella, Ph.D., S.E., P.E., F.ACI, F.ASCE, Michael Mota, Ph.D., P.E., SECB, F.ASCE, F.ACI, F.SEI, and Amy Trygestad, P.E., F.ACI

David A. Fanella is Senior Director of Engineering at the Concrete Reinforcing Steel Institute and can be reached at dfanella@crsi.org.

Michael Mota is Vice President of Engineering at the Concrete Reinforcing Steel Institute and can be reached at mmota@crsi.org.

Amy Trygestad is Director of Codes and Standards at the Concrete Reinforcing Steel Institute and can be reached at atrygestad@crsi.org.

layers in the concrete slab (Figure 1). Mild reinforcing bars are commonly used as the primary flexural reinforcement in the slab. Once the top layer of concrete is cast over the cage modules, a two-way slab system of uniform thickness is created.

Flat plate voided concrete slab systems are designed and detailed in accordance with Building Code Requirements for Structural Concrete (ACI 318) just like any other two-way slab system.

It is important to note that the void formers do not contribute to the nominal flexural and shear strengths of the slab system; their only role is to provide voids in the slab. More information on these systems can be found in CRSI publications, Design Guide for Voided Concrete Slabs and Frequently Asked Questions (FAQ) About Flat Plate Voided Concrete Slab Systems.

Construction – Part 1: General Requirements (ISO 834). ISO 834 and ASTM E119 time-temperature curves are also essentially the same, and it has been shown that the differences in severity between the two tests are negligible. Therefore, it follows that the results obtained from assemblies tested by the DIN requirements would be the same as those that would be obtained if the assemblies were tested in accordance with ASTM E119 requirements.

A fire test to supplement those performed previously was conducted in June of 2017 at the Fire Testing Laboratory of NGC Testing Services in Buffalo, NY. Information on the test assembly and methods, and a summary of the results, are presented below.

Test Assembly The test assembly consisted of an 8-inch thick concrete slab that was14 feet by 18 feet (the plan dimensions were limited by the size of the furnace test frame). The slab contained 4-inch-thick ellipsoidal void formers with a diameter of 123⁄8 inches that were spaced 13¾ inches on-center within the cage modules. A total of 140 void formers were used in the slab.

Normal-weight concrete with siliceous aggregate and a design compressive strength of 5,000 psi was specified. Slab reinforcement consisted of ASTM A615 Grade 60 #4 bars spaced 12 inches on-center in both primary directions at the top and bottom of the slab with a ¾-inch clear cover. The cage modules were tied to these reinforcing bars, and a 13-inch-wide strip of solid concrete was provided around the perimeter of the slab. Construction of the assembly followed typical construction procedures used in the field for these types of systems.

The layout of the void formers in the slab before casting of the concrete is shown in Figure 2.

The slab was cured for 28 days prior to testing, and its moisture content was measured to be 69% on the day it was tested.

An 8-inch-thick slab was tested because that is the minimum slab thickness that can be specified for

Fire Resistance Numerous fire tests have been performed on BubbleDeck systems and Cobiax systems in accordance with the provisions in Fire Behavior of Building Materials and Building Components; Building Components; Definitions, Requirements, and Tests (DIN 4102-02). Fire-resistance ratings of at least 2 hours were obtained from these tests for flat plate voided concrete slabs with a cover of ¾-inch to the main flexural reinforcing bars. The time-temperature curve used to test assemblies in the DIN requirements is essentially the same as that prescribed in Fire-resistance Tests – Elements of Building

Figure 1. Flat plate voided concrete slab system.

12 January 2018

Figure 2. Layout of void formers in test assembly

or rotation. As such, the assembly was unrestrained during the duration of the fire test, which is usually conservative for cast-inplace concrete slab systems.

The assembly was loaded during the entire time of the test with a uniformly distributed load of 80 psf using water-filled steel tanks. This load meets the ASTM E119 load requirement for floor and roof systems: “Throughout the fire-resistance test, apply a superimposed load to the test specimen to simulate a maximum-load condition. This load shall be the maximum-load condition allowed under nationally recognized structural design criteria unless limited design criteria are specified, and a corresponding reduced load is applied.”

The vertical deflection of the assembly was measured throughout the test with 3 plumb bobs located at the center and quarter points of the slab.

Results

Figure 3. The underside of test assembly after fire testing.

this type of system given the current product lines available. Because of savings in selfweight, long spans (and slabs thicker than 8 inches) are typical. For example, a 15.5-inchthick slab can easily span 40 feet and longer for residential and office loading. Applying the fire rating obtained from the 8-inch assembly to commonly used thicker slabs is conservative.

Method The assembly was tested in accordance with the requirements of ASTM E119. The furnace combustion chamber was fitted with 80 uniformly located natural gas burners, which provided an even heat flux across the assembly’s exposed surface. Furnace temperatures were maintained in accordance with the ASTM E119 time-temperature curve and were measured and recorded at 15-second intervals. The surface temperature of the unexposed surface was measured and recorded at 15-second intervals as well using 9 thermocouples located in accordance with ASTM E119 requirements.

The edges of the slab on all 4 sides were supported vertically by the test frame; no restraint was provided for thermal expansion

A few minutes into the fire test, sporadic spalling of concrete occurred beneath the void formers. Some of the void formers melted and dripped on to the furnace chamber floor while others remained mostly intact.

At approximately 30 minutes into the test, spalling decreased significantly as did the rate of vertical deflection of the assembly.

At 2 hours into the test, none of the ASTM E119 conditions to terminate testing were met.

At 2 hours and 52 minutes, the test was terminated because the temperature recorded at one of the thermocouples at the unexposed surface exceeded the ASTM E119 individual limiting temperature rise of 325 degrees Fahrenheit. The maximum deflection was approximately 3.5 inches at that time. Throughout the duration of the test, the assembly supported the applied loading with no signs of collapse.

The underside of the assembly after the fire test is shown in Figure 3. Although the concrete had spalled at various locations, the exposed reinforcing bars exhibited minor to no damage due to the fire.

Equivalent Slab Thickness and Fire Rating Table 722.2.2.1 of the International Building Code (IBC) provides minimum slab thickness of reinforced concrete floor and roof assemblies to achieve fire-resistance ratings based on aggregate type. Flat plate voided concrete slab systems are similar to slabs with ribbed or undulating soffits, so an equivalent slab thickness must be calculated for use in Table 722.2.2.1.

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In general, the equivalent thickness of a flat plate voided concrete system is equal to the net volume of concrete divided by the floor area.

The net volume of concrete is equal to the volume of concrete of a solid slab minus the average concrete displaced by the void formers: Volume of solid 8-inch slab = 8/12 = 0.67 cu ft/sq ft For the void formers used in the test assembly, average concrete displaced = 0.18 cu ft/sq ft Net volume of concrete = 0.67 – 0.18 = 0.49 cu ft/sq ft Thus, the equivalent thickness of the test assembly = 0.49 ft = 5.9 inches An equivalent thickness of 5.9 inches corresponds to a 2-hour fire-resistance rating for a normal-weight concrete mix with siliceous aggregate (according to Table 722.2.2.1, a 5-inch thickness is required for a 2-hour rating, and a 6.2-inch thickness is required for a 3-hour rating). The fire rating based on equivalent thickness is essentially the same as that determined from the fire test.

Conclusion Based on the results from the fire test, it can be concluded that the flat plate voided concrete assembly, constructed of the materials and in the manner described previously, achieved a 2-hour unrestrained assembly rating when exposed to fire in accordance with the test method prescribed in ASTM E119. This test confirms the appropriateness of the application of Appendix X3 of ASTM E119 and Section 703.3 of the International Building Code (ICC 2015) for flat plate voided concrete slabs.

This fire rating is conservative because, in actual structures, cast-in-place reinforced concrete floor and roof slabs subjected to fire are usually restrained by adjoining monolithic slabs and vertical elements. Table X3.1 in Appendix X3 of ASTM E119 provides guidance for determining restrained and unrestrained conditions of construction. According to that table, all types of concrete cast-in-place floor or roof construction, where the floor or roof construction is cast with the framing members, are considered to be restrained with respect to effects of thermal expansion.

No signs of structural collapse of the assembly were evident at any time during testing.

The testing was terminated because of the amount of heat transmission through the slab.

Thus, at least a 2-hour fire-resistance rating can be achieved for thicker flat plate voided concrete slabs.▪ The online version of this article contains references. Please visit www.STRUCTUREmag.org.

Practical SolutionS solutions for the practicing structural engineer

H

ave you ever been asked to provide an underlayment recommendation to level, resurface, or fill space over a structural element? This article can provide a framework for developing appropriate options.

By definition, underlayments are not structural.

The term underlayment is often used in reference to several types of products including poured toppings, either cement or gypsum, and sheet materials. This article focuses on underlayments defined as the layer between a structural subfloor and a finished floor that facilitates leveling and adhesion. Underlayments commonly enter the conversation when there is a specified floor height or when a correction of the structural floor element is required. Although both cement and gypsum topping products present self-leveling characteristics, many factors drive the decision of which product type best suits the project at hand. Armed with answers to a few project-related questions, narrowing that scope is a relatively straightforward process.

Self-Leveling Self-Help Key Differences and Practical Applications for Cement and Gypsum Underlayments

By Beth Lee and Josh Jonsson

Beth Lee (blee@maxxon.com) is the Business Development and Project Manager for Maxxon Corporation in Hamel, MN. Josh Jonsson is the Vice President of Sales for Maxxon Corporation in Hamel, MN and can be reached at jjonsson@maxxon.com.

What are you trying to accomplish?

To identify whether your project requires a gypsum or cement underlayment, the most important question to ask is, “What problem or need am I addressing?” Figure 1 provides several common floor issues mitigated by each type of product.

When asked to recommend a lightweight solution to meet a specific floor height, it is important to consider the depth of space to be filled. As a structural engineer, you might be asked to strike a balance between structural needs and weight considerations in a project where a specific thickness has been included in the project plans. Gypsum and cement underlayments possess lower densities than regular concrete while also achieving a flat surface ready for floor coverings. This makes them viable options for filling space and meeting thickness

requirements. In applications that require significant thickness to be filled, an alternative lightweight underlayment might be considered as an addition to the floor system. One such option is Expanded Polystyrene Foam (EPS), a highly cost-effective, rigid sheet of foam that is well suited as a deep-fill solution in spaces with a consistent depth to be filled. Another option is a deep-fill screed, a gypsum or cement underlayment that has been pre-blended with a lightweight aggregate, such as EPS beads, which is mixed with water on the job site and installed as a liquid. This lightweight deep-fill screed can weigh as little as 30 pounds per cubic foot (pcf ) and provides a suitable replacement for EPS foam sheets when density is of the utmost concern. There are a few scenarios when a deep-fill screed may be more advantageous than EPS foam.

• Significant levelness issues: for example, severe undulations in the subfloor take less time to correct with a cementitious, lightweight product than with a foam board.

• Requirement of a fully bonded floor system throughout the entire floor assembly: EPS foam is loosely laid over the subfloor while a cementitious, lightweight underlayment will bond to the subfloor.

• Depth of space excludes the use of EPS: EPS foam with holes or EPS foam topped with reinforcement requires a minimum 1-inch topping over the EPS. EPS without holes or not topped with reinforcement requires a minimum 1½-inch topping over the EPS. A deep-fill screed requires only a ¾-inch topping.

• If conduit or cable needs to be encapsulated within the buildup, then a cementitious, lightweight topping will be less time consuming, as foam board would need to be cut and placed around the conduit or cable.

It is important to note that both EPS foam and deep-fill screeds require a topping with a weight that is equal to or greater than 110 pcf before flooring installation (Figure 2). The gypsum or cement underlayment topping provides the density and

Figure 1. Common uses for gypsum and cement underlayments.

14 January 2018

Figure 2. In this project, a lightweight deep-fill screed was installed and allowed to dry overnight.

The next day, the hardened screed was topped with a gypsum underlayment to provide the density and compressive strength needed for point load dispersion and floor bonding compatibility.

compressive strength required for adequate point load dispersion, as well as meeting the flatness and bond compatibility requirements for floor coverings.

What is the structural building material?

Because of wood’s flexural properties, underlayments used over wood subfloors must possess the ability to resist cracking caused by movement. Cement underlayments can achieve high, concrete-like compressive strengths. With proper floor preparation, they achieve excellent bonding to concrete subfloors. (Many cement underlayments require shot blasting of the concrete to a Concrete Surface Profile of 3-5. Check with the underlayment manufacturer for proper floor preparation procedures.) Although this makes them ideal for many concrete applications, the stiffness associated with cement binders make them susceptible to cracking. This phenomenon is exacerbated by the elasticity of the subfloor.

Given the typical end-uses of light wood frame construction, a high compressive strength is not often needed. However, if the underlayment will be exposed to construction traffic for an extended period, or its end use requires exceptional strength, a high-end gypsum underlayment can achieve a compressive strength in the neighborhood of 4500 pounds per square inch (psi) and offers the potential for better crack resistance than its cement counterpart.

While wood frame projects using an underlayment will typically necessitate the use of a gypsum underlayment, concrete floor slabs require further consideration to determine which type of underlayment provides the optimum solution. If the subfloor is concrete, several additional questions will assist in selecting the best solution for the project.

What is the available time for repair?

Often the flooring installer discovers levelness or surface quality issues on a job site.

Typically, they will also be the ones to correct the issue. Such a discovery may require that the problem is remedied quickly.

Cement underlayments pre-mix the binder, aggregate, and additives in the bag, creating a highly homogeneous mixture which is combined with water at the job site. Due to the product’s inherent levelness-seeking properties and the characteristic compactness of job scope, installation of a cement self-leveler is not complicated once proper technique is learned. These products are readily available to flooring installers at home improvement and specialty stores.

Additionally, several cement underlayments include calcium aluminate as part of the composition. Calcium aluminate will give a self-leveler rapid strength formation, provide self-drying characteristics, and include a robust performance ability when exposed to occasional moisture or exterior

Figure 3. The floor of this train station needed significant work prior to repurposing. In this photo, extensive floor height variations can be seen in the locations where tracks were removed.

STRUCTURE magazine

elements during the construction process.

This makes it ideal for time-sensitive projects by significantly shortening the waiting period before the floor covering installation. These characteristics also allow it to be installed earlier in the construction process if needed.

Typically, gypsum underlayments are installed by mixing the gypsum binder, locally sourced sand, and water into a pump. The liquid screed is sent through a hose into the building where a crew installs the material to a specified depth. The crew runs periodic quality checks of the material, ensuring consistency throughout the project scope. Due to the size of most projects, gypsum underlayments are sourced directly from the manufacturer and often shipped directly to the job site. This requires that jobs be scheduled in advance to allow time for crew mobilization and shipping of the material. It should be noted that a limited selection of bagged, pre-sanded gypsum underlayments are available through retail outlets to meet the needs of uncommon specialty projects.

What is the budget?

Cement underlayments can be quite expensive due to the cost of the binder and additives.

This is offset to some degree by the inclusion of silica sand in the bag; however, the cost remains prohibitive for large spaces unless other factors demand their use. Large jobs requiring a cement self-leveler will typically be outsourced to a company that specializes in the pumped application of underlayments.

In large projects, a gypsum underlayment can provide a suitable, cost-effective alternative. High-end gypsum underlayments offer a quality compressive strength and the ability to be poured extremely thin (Figure 3 and Figure 4). If absolute flatness is required, two options exist:

Figure 4. The same area is pictured following the installation of a high-end gypsum underlayment, which leveled the floors and created a smooth surface ready for floor coverings.

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Figure 5. Gypsum and cement underlayment average material square foot price. Installed cost can vary widely depending on the scope of a project.

1) A cement underlayment can be used.

Cement underlayments contain exceptionally fine silica sand, which allows them to achieve a true featheredge.

2) A gypsum underlayment can be installed but will need to be buffed once it is dry.

Buffing is required because the aggregate added to a gypsum underlayment is coarser than silica sand.

Figure 5 is an average cost continuum of an economical gypsum underlayment, a highend gypsum underlayment, and a cement underlayment. These numbers are material cost averages; installed cost can vary widely depending on the scope of a given project.

As you can see, the cost between a standard underlayment and a cement underlayment can vary significantly. Your project conditions will dictate which product is best suited for the application but knowing that several options exist will pay off either through cost or time savings.

What is the depth of the levelness correction that is needed?

Cement underlayments are designed with the primary purpose of self-leveling. The composition of these underlayments include additives that make them more viscous. This compels the material to level out on its own, which makes them exceedingly effective in thin applications. Therefore, any application that requires a very thin layer of material (featheredge to < ¼ inch) will often justify a cement underlayment (Figure 6).

Floors with areas that are significantly outof-level will almost certainly benefit from a gypsum underlayment. They provide an economical solution, are available from 2000 psi to 4500 psi, and can be installed thick or thin. For depths greater than 1½ inches, the installer may recommend the underlayment be installed in two lifts to facilitate drying.

If the project is a renovation of a gypsum subfloor, resurfacing should always be performed using a gypsum underlayment or gypsum compatible patch to ensure a strong bond.

Unique project variables may come up during an underlayment discussion. These might include acoustical requirements, the presence of excessive moisture vapor in the concrete slab, or a project with strict subfloor requirements.

Some, though not all, uncommon project

variables may influence the type of underlayment that best suits the project.

Are there unique considerations for the project?

Both gypsum and cement underlayments may provide airborne sound reduction, which is measured as Sound Transmission Class (STC).

Airborne sound reduction is a function of mass; the greater the mass, the more significant the reduction. Given that gypsum underlayments are installed at a much greater depth than cement underlayments, it is reasonable to infer that a gypsum underlayment will give a better STC performance. In fact, gypsum underlayments are commonly used in tandem with sound control mats as part of a floating floor/ ceiling sound control system. The total system of a sound mat under a gypsum underlayment is installed to address both types of sound, airborne and impact sound waves. Installation of cement underlayment over a sound control mat is not recommended. The stiffness of the cement underlayment, combined with the thinness at which it is installed, can cause widespread cracking.

Gypsum and cement underlayments possess similar moisture vapor transmission characteristics. Neither type of underlayment will block the passage of moisture vapor; however, both can be installed over moisture mitigation systems, such as an epoxy vapor barrier. Like most cement underlayments, an epoxy vapor barrier requires floor preparation, such as shot

blasting or removal of surface contaminants, to ensure the epoxy is absorbed into the top of the concrete slab.

Though most floor coverings will only require around 2,000 to 3,000 psi for installation, thin glued flooring materials in commercial projects will require a high compressive strength.

Commonly used in hospitals, schools, and sports floors, thin resilient floor finishes are more likely to telegraph imperfections in the subfloor. A high-strength underlayment is more tolerant to the high point load needs of commercial floors. This type of project will likely include additional floor preparation on the front or back-end. This is because cement self-leveling underlayments that can achieve a zero edge require the floor to be shot blast prior to installation of the underlayment, while high-end gypsum underlayments will likely require rework (buffing) of the underlayment following installation. If a significant portion of the average floor depth is less than ¼-inch, it likely makes sense to invest in preparation work and a polymer modified cement-based self-leveling underlayment.

While gypsum and cement underlayments share many similarities, understanding their subtle differences will enable you to confidently recommend the best solution for any project. The rationale behind product usage often comes from different parties on a project. The owner of the project has a timeline and cost in mind, the general contractor is mindful of budget related to floor corrections, the structural engineer is alert to weight considerations, and the flooring installer is concerned with floor covering considerations. The key to making a wellinformed underlayment decision is to ensure all the details – assembly, end-use, timeline, and unique variables – are well understood and addressed.▪

Figure 6. Installation of a cement underlayment in which the thinness of the application can be observed. Prior to installation, the concrete subfloor had to be shot blast to ensure adequate bonding of the underlayment.

STRUCTURE magazine

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January 2018

New Member Benefits Including Unprecedented Digital Access to the Institute’s 200+ Guides and Reports

Other new benefits for individual members include:

• Free media mail shipping • Discounted access to ACI codes and specifications • Discounted access to ACI University’s monthly webinars and on-demand course subscription

Plus ACI’s existing membership benefits.

www.concrete.org/membership

Structural PracticeS practical knowledge beyond the textbook

P

art 1 of this article (STRUCTURE, October 2017) showed that if column and wall supports restrain the shortening of a post-tensioned member at stressing, part or all of the precompression intended for the post-tensioned member will be diverted to the supports. The loss of precompression to the supports will reduce the member’s moment capacity. This article outlines the mechanism for the development of moment capacity in fully restrained members. It also highlights the differences between members with bonded tendons and members with unbonded tendons with respect to moment capacity when subjected to support restraint. (Note: The Figure and Equation references continue the numbering from Part 1 of this series.)

Support Restraints and Strength of Post-Tensioned Members Part 2

By Bijan O. Aalami, Ph.D., S.E.

Bijan O. Aalami is Professor Emeritus at San Francisco State University and Principal of the ADAPT Corporation. He may be reached at bijan@adaptsoft.com.

Members Reinforced with Unbonded Tendons

B, since the tendon strand can slide within its sheathing. At the ultimate limit state, the restraint force from the supports increases to F4 (Figure 5c).

Under service conditions, represented by Figure 5a, the force at support A is F3 and the tendon force at the crack is F2. Since the crack extends through the depth of the member, F2 = F3. In Figure 5b, the elongation of the tendon increases its force at the crack from the service condition F2 to F, as shown in Figure 5c. The external force demand at the location of the crack is the axial tension N, moment M, and shear force V, shown in Figure 5d. The axial tension N equals F4, the restraint of the support at point A at the ultimate limit state, and from equilibrium of forces along the member N = F4.

The demand actions N and M at the cracked section are resisted by the increase in tendon force across the crack resulting from the displacement of the member and the compressive force (C ) developed at the newly established contact point.

The relationships are: C = F – F4 Equation 6 M = C z, Equation 7 M = (F –F4)z Equation 8 Where z is the lever arm between the centroids of the tension and compression forces, and F is the force in the tendon at the crack. Note that the tensile force in the tendon that contributes to the resistance capacity of the cracked section is the difference between the force in the tendon at the crack (F ) and the restraint of the support (F4).

Although the presence of non-prestressed reinforcement will help to disperse the cracks and control crack width, the contribution of the non-prestressed reinforcement to moment capacity is ignored to

Figure 5 shows a member reinforced with unbonded tendons with a single restraint crack that extends through its depth (Figure 5a). The crack resulted from shrinkage of the concrete and the restraint from the fixed end supports A and B. Supports C and D are assumed to provide no restraint, so are shown as roller supports. The tendons are shown as straight rather than profiled to simplify the presentation.

Note that, in Figure 5a, the tendon retains its force (F2) across the crack, but there is no compressive force across the face of the crack because the tendon force is diverted to the supports A and B. The force in the tendon at service (F2) is equal to the restraint force (F3), which for simplification is not shown in the figure.

An idealized free body diagram of the post-tensioning forces for the left segment of the member is shown in Figure 5c. With an increase in the applied load, the cracked sections will contact at the locations marked in Figure 5b and a compressive force C will develop at the point of contact. Also, the downward displacement of the slab before collapse will elongate the tendons, resulting in an increase (δF2) in the tendon force. The initial tendon force at the location of the crack under service condition (F2) will increase to its final value F as shown in Figure 5c. The increase in tendon force across the crack will be partially trans- Figure 5. Failure mechanism and partial force diagram of the ferred to the restraining supports A and member with restraint crack.

18 January 2018

Figure 7. Partial free body diagram; non-symmetrical cracking.

determined, the design capacity the section. The force (Fu – F4) is the friction of the section is known. Note that force between the strand and its sheathing at in an actual structure, the con- the ultimate limit state. tribution of the non-prestressed It is concluded that when members reinreinforcement across the crack forced with unbonded tendons are subject must be included; the compressive to significant support restraints, the friction force C will be resisted by both the between the steel strand and its sheathing Figure 6. Development of friction force P at ultimate strength. tendons and the non-prestressed plays a significant role in the strength available focus on the effects of support restraint, post- reinforcement. The capacity of the section from the tendon at the member’s ultimate tensioning, and moment capacity. depends on the location and the magnitude strength capacity. The larger the friction force, A free body diagram of the horizontal forces of the tendon force, and the location and the greater is the available tendon strength to for the left segment of the member of Figure amount of the non-prestressed reinforcement. resist applied loads.

5 is shown in Figure 6. Figure 6a shows the In the general case, a restraint crack is likely segment at impending collapse, where force to break the member into two non-equal Members Reinforced with C has developed at the contact point between lengths as illustrated in Figure 7. For static Bonded Tendons the two sides of the crack. The figure shows equilibrium of the member shown, the the contribution of the friction forces (P ) restraining forces (F4 ) on each side of the Members reinforced with bonded tendons will between a strand and its sheathing in devel- crack must be equal; likewise, the magnitude develop a larger moment capacity at restraint oping the compression force (C ). As shown of the friction force (P ) on each side of the cracks than members that are reinforced with below, the compressive force C that can crack must be the same. The friction force is the same amount of unbonded post-tensiondevelop across the crack before the collapse a function of several factors, including the ing. There are three reasons for this. of the member is limited to the friction force tendon length. The friction force that can be First, bonded tendons typically develop their (P) that builds up between a strand and its sustained across the crack will be that from specified strength (fpu) before failure, whereas sheathing (Figure 6b). the segment with the smaller force – typically members reinforced with unbonded tendons From Figure 6b: the shorter side of the member. Because C = tend to undergo large deflections and reach P = F–F4 Equation 9 P, the moment capacity is: failure due to crushing of concrete or excesFrom Figure 6a: M = Pz Equation 13 sive deflection before the tendons can reach C = F – F4 Equation 10 In summary, for the condiWhere F4 is the restraint from the support tions discussed, the maximum at the ultimate limit state. Therefore, tensile force that is available to C=P Equation 11 develop the resisting moment To arrive at the upper-bound for the at the crack is limited to the moment that can develop at the crack, the friction that develops between tendon is assumed to be stretched to rupture the tendon and its sheathing at before failure, even though this is unrealistic the ultimate limit state. This is for unbonded tendons. further illustrated in Figure 8, The force F in the tendon is calculated as: where Figure 8b shows the idealF = Aps fpu Equation 12 ized in-service and the ultimate Where Aps is the tendon cross-sectional area, limit state distribution of the and fpu is its specified strand strength (com- force in the tendon. In Figure monly 270 ksi; 1860 MPa). In practice, fpu 8b, the tendon force in service, is unlikely to develop for unbonded ten- F2, at the crack location, is equal dons; this is merely a hypothetical case for to the support restraint F3. At an upper-bound limit. The tendon force F the ultimate limit state, the will decrease along the tendon length due to tendon force at the crack locafriction between the tendon and its sheathing. tion increases to Fu. For a given tendon profile and friction coefIn the preceding diagram, the ficients, the stress loss due to friction can be force (Fu – F4) is the force that readily calculated [Aalami, 2014]. is available to resist applied Once the friction force P and hence the moments at the ultimate limit Figure 8. Service and strength limit force diagram of members compressive force C across the crack are state – the moment capacity of reinforced with unbonded tendons.

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Figure 9. Forces at ultimate limit state in a member with bonded tendon and restraint cracking.

Figure 10. Comparison of tendon force at service and ultimate limit (strength) states for bonded and unbonded post-tensioning tendons.

their specified strength. Consequently, the American Concrete Institute’s ACI 318 and Euro Code 2 (EC2) specify a significantly lower permissible stress (fps) for unbonded tendons when calculating flexural capacity.

ACI 318 limits the increase in tendon stress at ultimate strength for unbonded tendons to between 30 to 60 ksi (206 to 413 MPa), depending on the span to depth ratio. In EC2, the increase is limited to just 100 MPa (15 ksi). This is only a 7 to 9% gain in strength over service conditions, leaving about 30% of a tendon’s strength unused at member failure.

Second, for members with bonded tendons, the increase in demand moment at a point results in an increase in the

tendon force at that location. The local increase in tendon force is not offset by support restraint like it is for unbonded tendons.

Third, the fact that the friction between the strand and the duct during stressing of a bonded tendon is higher than the friction between the strand and the sheathing of an unbonded tendon is advantageous at the strength limit state of a cracked section.

Consider the member with a bonded tendon shown in Figure 9. Let the restraint from the supports be large enough to cause the cracking shown in Figure 9a. The force in the tendon at grouting follows the friction diagram shown in Figure 9b. Let the force in the tendon at the crack location under service conditions be F2. For static equilibrium, F2 must be equal to the restraint of the support (F3) while the gap at the crack is open. An increase in the applied moment at the crack location will tend to elongate the tendon locally, leading to an increase in the tendon force by δF2 (Figure 9c).

The demand actions at the crack location (Figure 9d ) are moment M, axial force N, and shear V. From equilibrium of the forces, N is equal to F3, the force due to restraint of the support.

The tensile force available to resist the demand actions at the crack location is: T = F2 + δF2 – F3 Equation 14 Since at the crack location F2 = F3, the available force (T ) to resist the induced moment is equal to δF2.

Likewise, from equilibrium of forces, the compression force C is C = (F2 + δF2 ) – F3 = δF2 Equation 15 The moment, M, that can be developed at the crack is equal to: M = Cz = δF2 z Equation 16

Comparison between Unbonded and Bonded Systems Figure 10 compares the effect of restraint cracking on the strength of a member with unbonded tendons to that of a member with bonded tendons. The diagram shows the distribution of force in the tendon for the segment to the left of the crack. Line ABC is the force in the tendon at service conditions.

The force is influenced by the friction between the tendon and its sheathing, or duct and the seating loss at stressing; to simplify the presentation, the force is assumed to be the

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same for bonded and unbonded systems at service conditions. For a bonded system, the ultimate moment at the crack location will include a local increase in the tendon force as shown by Point D (stress Fb). The tendon force available to resist the demand moment is equal to the local increase in the tendon force (δFb) shown by CD. For an unbonded tendon, the distribution of force at the strength limit is governed by the re-alignment of the friction force along the tendon length from line ABC to line EG. The force available to resist the moment demand is (F2 + δFu ) – F4 = Fu – F4.

Hence, the net force (T ) from the posttensioning tendons available to resist the demand moment at the crack is: Unbonded: T = Fu – F4 Equation 17 Bonded: T = Fb – F3 = δFb Equation 18

Conclusion This article explains the reduction in moment capacity of post-tensioned members due to the loss of precompression when restraint from wall and column supports prevents free shortening of the members during stressing. It concludes that the contribution of tensile force from prestressing tendons to the moment capacity of a post-tensioned member is tied to the value of precompression from post-tensioning. If the supports prevent all shortening, the entire post-tensioning force will be diverted to the supports, leaving the member with no precompression – hence, there is no contribution from posttensioning to moment capacity.

The article highlights the importance of detailing at the lowest levels of building structures, where the constraint of the supports is most pronounced. The addition of nonprestressed reinforcement helps to disperse the cracking and control the crack width but does not compensate for the loss in moment capacity caused by support restraint. The added non-prestressed reinforcement must not be less than the loss of contribution to moment capacity from post-tensioning due to support restraint.

The article also explains why the effect of support restraint is not the same for members reinforced with unbonded tendons and members reinforced with bonded tendons. It concludes that the loss of moment capacity due to support restraint is significantly greater for members with unbonded tendons than for members with bonded tendons.▪

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Structural Performance performance issues relative to extreme events

S

tructural engineers are not traditionally involved in the analysis or design of building fire safety. When they are, their focus is generally on structural fire protection and, with some exceptions, their scope is limited to ensuring compliance with prescriptive building code requirements for the fire resistance ratings of different building elements.

But structural fire protection is just one aspect of a comprehensive framework for building fire safety. As demonstrated by the Grenfell Tower fire in London, shown in Figure 1, structural fire protection alone does not ensure the fire safety of a building or its occupants. While structural engineers may not practice in the field of fire protection engineering, it is useful for them to have at least a basic understanding of building fire safety issues.

An Overview of Fire Protection for Structural Engineers Part 1

By Frederick W. Mowrer, Ph.D. and Richard L. Emberley

Frederick W. Mowrer is the Founding Director of Fire Protection Engineering Programs at Cal Poly in San Luis Obispo, CA. Dr. Mowrer is a Fellow of the Society of Fire Protection Engineers and a past-president of the Society.

He may be reached at fmowrer@calpoly.edu.

Richard L. Emberley is an Assistant Professor in the Mechanical Engineering Department and Fire Protection Engineering Program at California Polytechnic State University (Cal Poly). He may be reached at remberle@calpoly.edu.

This is the first of a 3-part series presenting an overview of fire protection for structural engineers.

This article addresses the fire safety objectives that drive the design of fire protection systems and features in buildings. The second article will address the different fire safety features and systems that are deployed to meet these objectives.

The third article will provide an overview of the emerging practice of structural fire engineering, which requires close collaboration between fire protection engineers and structural engineers.

the stakeholder-agreed objectives and criteria.

Performance-based design is most commonly used for complex buildings to demonstrate that the proposed design is at least “equivalent” to the prescriptive regulatory requirements for fire safety. This calls for an understanding of the implicit performance objectives associated with the regulatory requirements.

Life Safety For life safety analysis, the fundamental fire safety objective is to ensure that building occupants are not injured by a fire. From a performance standpoint, this objective is typically addressed by either entirely evacuating a building or relocating occupants to safe areas within the building.

Total evacuation is typically used for relatively small buildings, while partial or phased evacuation is more common for very large buildings, particularly high-rise structures. The fundamental objective associated with life safety from fire can be expressed as: Available safe egress time (ASET) > Required safe egress time (RSET).

The available safe egress time (ASET) depends on the rate of fire hazard development and the movement of smoke and heat throughout a building. In turn, the rate of fire hazard development depends on the flammability characteristics and configuration of fuels contributing to the fire, and on the presence or absence of automatic fire suppression systems to control or suppress fire development.

Fire Safety Before considering the different fire safety systems and features installed in buildings, it is useful to consider the fire safety objectives addressed by these systems, generally summarized as: • Life safety – for both occupants and emergency responders • Property protection • Mission continuity • Preservation of cultural heritage • Protection of the environment (from fire effluent and fire protection agents) Traditional prescriptive building regulations address these performance objectives implicitly by specifying the building fire safety systems and features required in different buildings. In performancebased fire safety design, performance objectives and criteria are explicitly stated, and engineering analyses are performed to demonstrate achievement of Figure 1. The Grenfell Tower (London) fire. Courtesy of Getty.

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The movements of smoke and heat from a fire depend on the compartmentation features of a building, including on individual floors, and for vertical pathways such as shafts, as well as on the response of mechanical ventilation (HVAC) systems under fire conditions. HVAC systems may be designed to simply shut down under fire conditions to prevent active circulation of smoke, or they may be designed to actively counter smoke flow, e.g., in pressurized exit stairs.

The required safe egress time (RSET) depends on many factors, including how quickly a fire is detected, how quickly occupants are notified, how effective the notification is (e.g., general evacuation alarm vs. voice communication system), how long occupants take to start the evacuation process, and, finally, how long it takes occupants to reach a place of safety once they start moving.

Property Protection

Mission Continuity The objective of mission continuity addresses potential disruptions to the mission of an enterprise or operation caused by fire. In the commercial insurance world, this type of loss is sometimes referred to as “business interruption.”

The financial magnitude of such losses

can far exceed the magnitude of direct property losses for some operations. However, these losses may be of primary concern to the mission operator and its insurance interests rather than to society in general.

Mission continuity can also be important at a community level. For example, a major fire that shuts down a hospital can have an impact on an entire region. A fire in a factory that employs a large percentage of

the local population can also have a widespread impact on the local economy. In a performance-based design framework, such impacts may be addressed by the stakeholders, but in a prescriptive design framework, such impacts may not be addressed.

The mission continuity objective is typically addressed with the same fire protection systems and features used to meet the property protection objective. Both property

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The objective of property protection is to limit the extent of physical damages to a level acceptable to the various stakeholders on a project. In general, the objective of property protection is of primary interest to property owners and their insurers; it is generally of less interest to building regulators, so it is not addressed explicitly in building regulations. However, property protection is addressed implicitly in building regulations in terms of height and area limits on buildings based on construction type, and in terms of requirements for automatic sprinkler protection.

The objective of property protection is commonly addressed through the application of automatic fire suppression systems. Automatic sprinkler systems are the predominant choice for property protection, but “clean agent” systems are frequently used in applications where water and smoke can damage sensitive equipment, such as electronics. Property protection is also addressed through compartmentation with fire barriers separating different fire areas to limit the maximum expected size of a fire and the maximum amount of property exposed to a fire.

Figure 2. Example event tree.

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Seattle Tacoma Lacey Portland Eugene Sacramento

San Francisco Los Angeles Long Beach Pasadena Irvine San Diego

Boise St. Louis Chicago Louisville New York

1 minute after ignition.

2 minutes after ignition.

3 minutes after ignition.

3 minutes, 30 seconds after ignition.

Figure 3. Living room fire transitioning through “flashover.” Courtesy of Kerber, S., “Impact of Ventilation on Fire Behavior in Legacy and Contemporary Residential Construction,” UL Firefighter Safety Research Institute, December 14, 2010.

protection and mission continuity objectives can be addressed by providing automatic fire suppression systems and by limiting the size of fire areas with fire barriers.

Preservation of Cultural Heritage The objective of preserving cultural heritage artifacts and structures has limited application but is recognized as distinct for at least two reasons. First, the artifacts and structures being preserved are irreplaceable; if they are damaged by fire, smoke, or water, they may lose their intrinsic value. The second reason, which applies primarily to structures, is because the historical fabric can be compromised if modern fire protection features are introduced without due consideration of their impact. Nonetheless, this objective is typically addressed through the use of fire detection and fire suppression systems to detect and control a fire before it causes unacceptable damage.

Environmental Protection Much like the previous objective, the environmental protection objective has had relatively limited application but is still worth mentioning. The primary ways to limit the potential environmental impacts of fire are to prevent fires and to control fires while they are still small. Where hazardous materials are involved, it may be necessary to use impoundment

structures to limit runoff of hazardous materials and contaminated firefighting water. It is also necessary to consider environmental impacts of firefighting agents, such as firefighting foams and halogenated agents, before they are specified for an application.

Defense-in-Depth Approach to Fire Safety Design The fire safety objectives discussed above differ from each other in terms of their goals, but the methods and means of achieving these objectives are often similar. The concept of defense-in-depth is fundamental to building fire safety design. As it applies to fire development, this concept can be summarized as: • Prevent fires from igniting; • Detect fires that do ignite while they are still small and manageable; • Suppress fires while they are still small and manageable; and • Confine fires that are not suppressed early with fire-resistive construction.

Different fire protection systems and features are designed to interact with and mitigate fires at different stages of their development.

These systems and features are discussed in the second article in this series. Because the fire safety interventions occur sequentially, the probabilities and consequences associated with fire are sometimes represented using

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event trees, which show how the consequences change depending on the success or failure of each sequential mitigation strategy. An example fire event tree, with the associated fire safety features and systems that intervene at each stage, is shown in Figure 2, page 23.

One of the distinguishing aspects of fire safety design is that the “load” imposed by a fire changes dynamically as the fire grows, so the sooner successful intervention occurs, the smaller the impact will be. This is illustrated graphically in Figure 3, which shows the rapid development of a contemporary living room fire to “flashover” or “fully developed” conditions.

Because of the threat flashover presents with respect to all the fire safety objectives discussed above, one common fire safety strategy is to prevent flashover, typically through the installation of automatic sprinkler systems.

However, automatic sprinkler protection is not effective for all fire scenarios, so other mitigation measures still need to be considered. In particular, structural fire protection becomes important once a fire reaches flashover conditions, both in terms of confining the fire with fire barriers as well as maintaining the structural integrity of the building.

The defense-in-depth design approach used for fire is different from other building design disciplines, which frequently design for maximum or steady-state load conditions rather than transitory conditions.

Summary The design of building fire safety typically uses a defense-in-depth approach to achieve fire safety objectives. In a prescriptive environment, the building code specifies the fire safety features and systems required to achieve the implicit fire safety objectives. In a performance-based environment, designers specify the fire safety features and systems needed to achieve stakeholder-agreed performance objectives. These fire safety features and systems will be discussed in the next article in this series.

The traditional role of the structural engineer in building fire safety design has been limited. However, this has been changing over the past two decades with the emergence of structural fire engineering as a distinct design discipline. This new role for structural engineers in the design of building fire safety will be discussed in more detail in the third article in this series.▪ The online version of this article contains references. Please visit www.STRUCTUREmag.org.

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Structural Systems discussion and advances related to structural and component systems

T

he lateral force-resisting system tends to be one of the most challenging aspects of the structural design of a building.

In todayâ&#x20AC;&#x2122;s wood-framed construction, designers consistently see larger buildings combined with bigger and more numerous window and door openings. This construction trend usually translates into reduced areas for lateral resistance throughout the structure.

There are several ways to address this challenge. The American Wood Council (AWC) publication, Special Design Provisions for Wind & Seismic (SDPWS), provides designers three acceptable methods for designing wood shear walls to resist lateral forces: 1) Individual FullHeight Wall Segments, the â&#x20AC;&#x153;traditionalâ&#x20AC;? design approach; 2) Perforated Shear Walls, an empirical design method based on the percentage of full-height wall segments adjacent to openings; and 3) Force Transfer Shear Walls, referred to in this article as force transfer around openings (FTAO), allows the utilization of the full wall geometry, including sheathed areas above and below openings.

Wood-Framed Shear Walls Utilizing Force Transfer Around Openings

By Jared S. Hensley, P.E.

Jared Hensley oversees the Pacific Northwest territory for APA The Engineered Wood Association.

He may be reached at jared.hensley@apawood.org.

Historically, the FTAO method has been considered a last resort in the design industry, due in large part to the fact that it is more comprehensive and therefore assumed to be more time-consuming than the other two design options. The application of the FTAO method also varies greatly within the industry, as the code merely states that the design must be based on rational analysis. With no clear directive in the design codes, there have been multiple design rationale techniques developed (e.g., Drag-Strut Analogy, Cantilever Beam, Diekmann), which has caused some dispute between design professionals as to which technique is the most precise. These discrepancies are what led APA to perform full-scale wall tests on the FTAO design methodology, intended

to prove that the application of FTAO can be more practical and simultaneously provide the most design flexibility.

This article discusses the benefits of FTAO for wood-framed structures, illustrates the testing performed on wood structural panel sheathed wood-framed shear walls for the enhancement of FTAO methodology, and provides a design example for a wood-framed shear wall with multiple openings and asymmetric wall piers.

Benefits of FTAO Wood Shear Walls There are many benefits to the FTAO method, such as utilizing the entire shear wall geometry to resist the applied load, potentially reducing the number of hold-downs, and reducing the base plate shear anchorage. However, the most significant benefit of FTAO over other shear wall design methods lies within the definition of the wall pier aspect ratio.

In the segmented and perforated methods, the aspect ratio of the contributing wall piers is defined as the full height of the wall system divided by the length of the wall pier (Figure 1a and 1b). When utilizing the FTAO method, the aspect ratio of the wall pier is defined as the height of the opening divided by the length of the adjacent wall pier (Figure 1c).

For example, given the wall illustrations in Figure 1 where the full height (h) is equal to 10 feet and the opening is 4 feet tall, the minimum wall pier length (b) allowed to be included in the design at the 2:1 aspect ratio for each method would be: segmented and perforated are equal to (10/2) or 5 feet and FTAO is equal to (4/2) or 2 feet. This benefit not only allows for narrower wall segments but also helps increase the overall height of the wall system where other methods might be limited.

This benefit is furthered by utilizing wood structural panel sheathing. The governing aspect ratio limitation in most applicable building codes for wood structural panels is 2:1, with a maximum of 3.5:1 (with the application of an adjustment factor). Although the use of other building

Figure 1. Illustration of aspect ratio (h/b) comparison for all three shear wall methods.

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Figure 2. Vertical shear line illustration.

Figure 3. FTAO deflection illustration.

materials is permitted, the aspect ratios either cannot exceed 2:1 or a load reduction factor is applied when the aspect ratio surpasses 1:1 per SDPWS Section 4.3.4.

Testing FTAO Design Methodology

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FTAO has become more popular amongst design professionals as a viable shear wall analysis for wood-framed walls. The majority of the published FTAO design examples, however, are limited to the wall containing a single opening and, until recently, the recognized FTAO methods were tested in a limited number of installations.

In 2009, a joint research project was conducted by APA–The Engineered Wood Association (APA), the University of British Columbia (UBC), and the USDA Forest Products Laboratory to examine the internal forces generated in wood structural panel sheathed wood-framed shear walls during a lateral event. The test results, in conjunction with analytical computer-based models from UBC (twelve full-scale, 8-foot x 12-foot wall configurations) were used to develop an enhanced FTAO design methodology and evaluate the accuracy of the calculated forces in the walls using historic FTAO methods.

The initial testing determined the Diekmann rational analysis as the most accurate simplified method to estimate the forces in the shear wall. Using this approach as the basis, APA expanded the methodology to incorporate multiple openings and asymmetric piers. This methodology is based on the following key design assumptions: 1) The unit shear is equivalent above and below the openings.

2) The corner forces are based on the unit shear above and below the openings and only the wall piers adjacent to that unique opening.

3) The tributary length of the opening is the basis for calculating the shear to each wall pier.

4) The shear to each wall pier is the total shear divided by the length of the wall, multiplied by the sum of the pier length plus the tributary width of the adjacent openings, divided by the pier length: v1 = (V/L)*((L1 + T1)/L1) Equation 1

5) The unit shear in the corner zones is equal wall pier widths and a single opening in the wall to subtracting the corner forces from the system. Figure 4, page 28, illustrates the benefits panel resistance, R, which is equal to the of FTAO for walls with multiple openings and shear of the pier multiplied by pier length: asymmetric wall pier widths.

Va1 = (v1L1–F1)/L1 Equation 2 Given a 20-foot-long wall that is 10 feet tall 6) The design is checked by summing the with a 4,000-pound shear force applied to shears along each vertical line. The first and the top of the wall, use the FTAO method last lines sum to the hold-down force and to calculate hold-down forces, required strap the interior lines sum to zero (Figure 2). forces, and required wall sheathing capacity.

Testing data was also used to analyze the Begin by calculating the hold-down forces overall deflection of the FTAO wall systems, at each end of the shear wall: which verified that the sheathing below the Step 1. H = (V*h)/L = ((4,000 lb)(10 ft))/ openings aided in resisting the overall deflec- (20 ft) = 2,000 pounds tion of the wall. To attain the most accurate The selected hold-downs must have a fastener deformation variable, the basis of minimum capacity of 2,000 pounds. the testing was completed using the fourBoth the unit shear and boundary forces term deflection equation, 2015 International above and below the openings must be deterBuilding Code (IBC) Equation 23-2: mined to attain the corner forces. Δ = 8vh3/EAb + vh/Gt + 0.75hen + da h/b Step 2. Unit Shear Above and Below Openings: Note that the 3-term deflection equation va = vb = H/(ha + hb) = 2,000 lb/(2 ft + 3 provided in SDPWS (Equation 4.3-1) may ft) = 400 plf also be used, but the deflection calculations Step 3. Total boundary force above and must be consistent throughout the design. below the openings: The wall deflection assumption (Figure O1 = (va)(LO1) = (400 plf)(4 ft) = 1,600 pounds 3) is that the total deflection of the FTAO O2 = (va)(LO2) = (400 plf)(6 ft) = 2,400 pounds shear wall is equivalent to the average of Calculate the corner forces at each openthe deflection of each wall pier in both the ing to determine the maximum strap positive and negative directions. The wall force required: pier heights also vary depending on the Step 4. F1 = ((O1)(L1))/(L1 + L2) = ((1,600 deflection direction and amount of sheath- lb)(2 ft))/(2 ft + 4.5 ft) = 492 pounds ing below the openings. For example, in F2 = ((O1)(L2))/(L1 + L2) = ((1,600 lb)(4.5 Figure 3, positive deflection of wall pier 1 ft))/(2 ft + 4.5 ft) = 1,108 pounds (δ1+) is determined using the height mea- F3 = ((O2)(L2))/(L2+L3) = ((2,400 lb)(4.5 sured from the bottom of the opening to ft))/(4.5 ft + 3.5 ft) = 1,350 pounds the top of the wall due to the resistance F4 = ((O2)(L3))/(L2 + L3) = ((2,400 lb)(3.5 of the wall sheathing below the opening. ft))/(4.5 ft + 3.5 ft) = 1,050 pounds Negative deflection of wall pier 1 (δ1-) is determined using the full wall height.

Demos at www.struware.com Δ = average (δ1+, Wind, Seismic, Snow, etc. Struware’s Code Search program calculates these and δ2+, δ3+, δ1-, δ2-, other loadings for all codes based on the IBC or ASCE7 in just minutes (see online δ3-) Equation 3

Figure 4. Design example illustration.

Figure 5. Wall force diagram illustration.

The selected strapping to be placed above and below the openings must have a minimum tension capacity of 1,350 pounds.

To calculate the unit shear beside the openings and to determine the wall sheathing requirement, the tributary width of each internal shear line must be determined.

Step 5. T1 = ((L1)(LO1))/(L1 + L2) = ((2 ft) (4 ft))/(2 ft + 4.5 ft) = 1.23 feet T2 = ((L2)(LO1))/(L1 + L2) = ((4.5 ft)(4 ft))/ (2 ft + 4.5 ft) = 2.77 feet T3 = ((L2)(LO2))/(L2 + L3) = ((4.5 ft)(6 ft))/ (4.5 ft + 3.5 ft ) = 3.38 feet T4 = ((L3)(LO2))/(L2 + L3) = ((3.5 ft)(6 ft))/ (4.5 ft + 3.5 ft ) = 2.63 feet Using the tributary width to the internal shear lines, the maximum unit shear beside the openings can be calculated to determine the required wall sheathing capacity.

Step 6. V1 = ((V/L)(L1 + T1))/L1 = ((4,000 lb/20 ft)(2 ft + 1.23 ft))/(2 ft) = 323 plf V2 = ((V/L)(T2 + L2 + T3)/L2 = ((4,000 lb/20 ft)(2.77 ft + 4.5 ft + 3.38 ft))/(4.5 ft) = 473 plf V3 = ((V/L)(T4 + L3)/L3 = ((4,000 lb/20 ft) (2.63 ft + 3.5 ft))/(3.5 ft) = 350 plf The selected wall sheathing must have a minimum shear capacity of 473 plf.

Note that, as a calculation verification, the sum of the unit shears multiplied by the length of the adjacent wall pier should be equal to the initial shear force: (V1)(L1) + (V2)(L2) + (V3)(L3) = V (323 plf )(2 ft) + (473 plf )(4.5 ft) + (350 plf )(3.5 ft) = 4,000 lb (This checks.)

The final design assumption must be met to verify the calculations. The sum of the shear-line forces at the outside lines must be equal to the hold-down force, and the sum of each interior shear line must be equal to zero.

The corner force resistance and corner zone unit shears must be determined to perform this verification.

The corner force resistance, R, for each pier is determined by multiplying the unit shear in each wall pier by the corresponding wall pier length: Step 7. R1 = (V1)(L1) = (323 plf )(2 ft) = 646 pounds R2 = (V2)(L2)= (473 plf)(4.5 ft) = 2,129 pounds R3 = (V3)(L3)= (350 plf)(3.5 ft) = 1,225 pounds To calculate the corner force unit shear in each wall pier, the resultant shear force in the corner zone must be determined, then divided by the corresponding wall pier length: Step 8. R1-F1 = 646 lb – 492 lb = 154 pounds R2-F2-F3 = 2,129 lb – 1,108 lb – 1,350 lb = -329 pounds R3-F4 = 1,225 lb – 1,050 lb = 175 pounds Step 9. vc1 = (R1-F1)/L1 = (154 lb)/2 ft = 77 plf vc2 = (R2 – F2 – F3)/L2= (-329 lb)/4.5 ft = -73 plf vc3 = (R3 – F4)/L3 = (175 lb)/3.5 ft = 50 plf Performing a design check by summing the forces at each shear line completes the analysis: Step 10. Line 1: (va1)(ha + hb) + (V1)(ho) = H? (77 plf )(2 ft + 3 ft) + (323 plf )(5 ft) = 2,000 lb Line 2: (va)(ha + hb)–(va1)(ha + hb)–(V1) (ho) = 0?

(400 plf )(2 ft + 3 ft)–(77 plf ) (2 ft + 3 ft)–(323 plf )(5 ft) = 0 Line 3: (va2)(ha + hb) + (V2)(ho)–(va) (ha + hb)= 0?

(-73 plf )(2 ft + 3 ft) + (473 plf ) (5 ft)–(400 plf )(2 ft + 3 ft) = 0 Line 4 = Line 3 Line 5: (va)(ha + hb)–(va3)(ha + hb)–(V3) (ho)= 0?

(400 plf )(2 ft + 3 ft)–(50 plf ) (2 ft + 3 ft)–(350 plf )(5 ft) = 0 Line 6: (va3)(ha + hb) + (V3)(ho) = H? (50 plf )(2 ft + 3 ft) + (350 plf )(5 ft) = 2,000 lb

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Using the 4-term deflection equation (Figure 3; Equation 3), overall wall deflection is calculated by determining the average of the positive and negative deflection values for each pier (Figure 4). Assuming a maximum hold-down capacity of 2,500 pounds, nail slip of 0.125 inches, and an APA Rated Sheathing 15 ⁄32-inch performance category with 8d nails at 4 inches on-center: Step 11. Δ = average (δ1+, δ2+, δ3+,δ1, δ2-, δ3- = average (1.393,0.692,0.508,0.708,0.6 92,0.963) (in) = 0.826 inches APA has automated the calculation process in the previous example with the creation of an FTAO spreadsheet (www.apawood.org/designerscircle-ftao).

The spreadsheet is based on the tested methodology and provides engineers with the required hold-down forces, tension strap forces, and wall sheathing capacity.

Finally, the tool automatically completes the design check in Step 9. The spreadsheet also provides shear wall deflection calculations for the 3- and 4-term deflection equation options.

Conclusion With the introduction of a new method for estimating the overall wall deflection and tested verification for the accuracy of force transfer around openings (FTAO) shear walls, designers can be confident in having another valuable tool to apply when designing the lateral resisting system for wood buildings. The FTAO method helps expand the boundaries of building design with wood-framed walls, utilizing the full shear wall geometry to help increase the design flexibility and potentially reduce the economic impact of the lateral resisting system.▪

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CERTIFIED

Mountain “S” Home

By James M. Williams, P.E., C.E., S.E., AIA, LEED AP

S

tepping and twisting around its’ site, this “S” shaped, $25 irregularity. Each pier typically supports two separate roofs at two million mountain home provides 360-degrees of scenic views. different elevations due to the stepped nature of the structure. The Located above Park City, Utah, the Mountain “S” Home is roofs help brace the piers in their weak direction. a structural engineering Opus. Every design element of the Although the building supports heavy snow loads and must resist large home represents an engineering challenge. The Mountain “S” Home seismic forces, the perimeter walls are almost entirely glass. The majority was a winner of the 2016 NCSEA Excellence in Structural Engineering of the structure’s front, back, and end walls are floor to ceiling glazing, Awards Program, category – New Buildings under $10M. maximizing views and providing an abundance of natural light. In many The architect, Wallace Cunningham, had strict requirements when locations, you can look front to back, through the house, without any it came to the structure; the structure would not interfere with the obstructions. Many of the roofs have no visible means of support. Headers views, must fit within the curved floor plan and within the wall, roof, were not required above the large glass openings since each roof typically and eave thicknesses allotted, and would complement the architecture. spans parallel to the window glazing. Member deflection had to be limited It would also need to withstand to not over-stress the large glass heavy snow loads, large seismic openings. The wood LVL roof forces, and the harsh mountain joists are supported at each end environment. Using materials that by steel channels, in-plane with did not require much fabrication each roof. The steel channels are and that could be modified in the welded to embedment plates that field would also be required, due are cast into the concrete piers. to the short building season.

Each end of the channels is typiThe Mountain “S” home has cally cantilevered, some as much as 30 individual kite-shaped roofs, 21 feet. The large cantilevers could which, from above, resemble not show any noticeable deflecscales on a dragon. Each of the tion, even when subjected to the upper roofs appears to be supsignificant snow load. ported by clerestory windows on Foundations, suspended floors, all sides. The roofs are only 16 and piers are all constructed of inches thick and cantilever hori- The “S” shaped home steps and twists around its’ mountain site. Although the home is cast-in-place reinforced concrete, zontally as much as 21 feet while curved, all of the structural members are straight. while the roofs are primarily LVL supporting 235 pounds per square foot (psf ) of snow (the equivalent lumber with steel channels at each side. Although the floor plan is weight of 6 cars stacked over the entire roof ). The architect insisted curved and stepped, all of the structural members are straight and that the fascia not exceed 1 inch in thickness (or height) in order to repetitive to help minimize costs and simplify construction. have a “razors” edge, and that it not show any sign of deflection, even The open-riser stairs are an architectural and structural design feature while supporting the 235 psf snow load plus the additional weight and focal point. The stairs have treads supported by steel plates which of snow drifting and icicles. This required transitioning from wood are supported on steel bars that are doweled into the wall and cantileframing to a special stainless-steel fascia fin. ver the full tread width. One of the stair runs extends past an exterior The majority of the roofs are supported by 16-foot-tall concrete piers glass wall with no attachment and only one stepped-support stringer, that range from 2 feet to 8 feet in length and are only 12 inches thick. constructed of segmented-stepped-steel bar stock welded together.

These piers not only support the heavy gravity loads but also resist the There is additional plate steel around the stair opening to achieve a thin lateral loads due to wind and seismic forces, in each direction. The profile. Adjacent to the stairs is a glass elevator with minimal structure.

16-foot-tall piers are cantilevered vertically from the structure below, There are a couple of massive interior concrete chimneys, fireplaces, which in some cases is a suspended floor, resulting in a structural and hearths, all of which appear to float above the floor. These elements STRUCTURE magazine

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The lower cantilevered roofs span perpendicular to 6-inch-thick concrete walls, resulting in tension or compression loads of 80,190 pounds on each end of each wall.

Stair treads are supported by a single “saw-tooth” stringer, cut from bar stock and then welded back together. Other treads cantilever off of a concrete wall.

are actually cantilevered off of the ends of concrete shear walls. In addi- had already set-off prior to the concrete even being placed in the forms. tion to “floating” fireplaces, there is also a large vanity in the master The concrete supplier took responsibility and paid for the repair and bathroom that “floats” across the exterior windows. replacement of the concrete. The superplasticizer was then added to the The house has a large exterior patio area that is located over the concrete on-site, and the thin walls were successfully re-constructed. living space. There was a concern about waterproofing due to the Most of the main floor is constructed of suspended reinforced concrete amount of snow and severe weather in the area. slabs and beams, with a maximum span of 45 feet, The decision was to provide redundancy in the allowing the lower level driveway to extend through waterproofing by designing a concrete structure the house. The long spanning floor system also supthat would have “self-healing” cracks. The large ports wall piers which support the roof, creating exterior patios have a reinforcing ratio similar to structural irregularities that had to be addressed. that of a water tank. The increased reinforcing Multiple radial grid systems were required to lay results in concrete that is water-tight, thus proout the structure. Although the house is curved in viding additional water-proofing and protection multiple directions, all of the structural members for the living space below. A membrane was also are straight and repetitive. Often the best structures provided between the structural slab and a topping are the ones that can be simplified, and once simplislab. Additionally, decorative water features, also fied, repeated, thus reducing costs and speeding up constructed of reinforced concrete, were designed construction. Such was the case for this project. and reinforced for water-tightness.

This home exceeds the owners’ expectations and In addition to the concrete piers which support the expectation of all who visit it. It is one of a most of the roofs, there are 9 staggered wall segkind. Due to the complex nature of the structure, ments that range from 10 feet to 13 feet tall and the local building official relied on the structural are only 6 inches thick (a constraint also imposed engineer for all structurally related inspections (in by the architect). These thin wall segments supaddition to the special inspectors). port lower roofs which span perpendicular to The structure is suitable for its environment; it the 6-inch-thick concrete segmented walls and Roofs cantilever inside and outside, with appears to be open, light, and airy. The home is which are staggered 2 feet, resulting in a tension clerestory glass between. The clerestory glazing located high in the mountains. The large extesits between two independent roof structures. or compression load of 80,000 pounds on each rior glass walls provide unobstructed views in end of each wall. The connection required an embedment knife plate all directions. All of the structural materials and members were with (8) #8 dowels, 62 inches long. The 6-inch-thick pier has (10) #9 readily available, and easy to work with. The winding hallways, tall vertical bars in a 2-foot-wide jamb. This occurs in 24 locations. The interior walls, stepped roofs, and transom glass feels as if you are 6-inch-thick segmented wall sections were required by the architect hiking through some canyon narrows as you pass through the house. to allow for a thin stone veneer and glazing between the staggered Roofs cantilever inside and outside, with clerestory glass between. walls. A superplasticizer was required along with several mockups and The clerestory glazing sits between two independent on-site testing. Initially, the architect wanted these walls to be only 4 roof structures, so cantilever deflections had to be inches thick. The engineer requested at least an 8-inch thickness but limited. With the large expanses of glass, the structure was only granted 6 inches maximum. is hardly noticeable.▪ After several mockups, the thin walls were constructed. When the James M. Williams is President of AE URBIA aka J.M. Williams forms were removed, there were huge voids throughout the walls, and Associates, Inc., a Utah based architectural and structural which raised concern. The voids were so excessive that the walls engineering firm. He has served as the President of SEAU and could not be used. Concrete had to be jackhammered and removed has also served on the Board and Executive Board of Directors for while maintaining the reinforcing steel. The voids were the result of the TCA for 7 years. James is currently on the AIA’s Codes and a superplasticizer being installed too early, and the concrete trucks Standards Committee. taking too long to arrive at the site on a hot day. The superplasticizer STRUCTURE magazine

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VIRTUAL OUTRIGGERS and CREATIVE ENGINEERING Reaching New Heights in Mexico

By Chris Crilly, P.E., Mark Tamaro, P.E., and Roberto Stark, Ph.D.

Final at night. Courtesy of TIF Digital.

G

rowing urban sprawl in Mexico’s most significant metropolitan areas negatively impacts the way people work and live, and many are seeking homes closer to work and public amenities. As a result, the demand for tall buildings is increasing in several urban areas, including Monterrey, the country’s second-largest city. Thornton Tomasetti, Inc. collaborated with local firm Stark + Ortiz, S.C., to provide structural engineering for the 279-meter (915-ft) Torre Koi, the tallest building in Mexico and third-tallest in Latin America.

Located just 5 miles (7 kilometers) from the city center, the 69-story Torre Koi is the centerpiece of the VAO complex in San Pedro Garza García, an inner suburb of Monterrey. The slender, mixed-use tower comprises 20 levels of office space, 218 apartments and 18 penthouses on 35 residential levels, and an amenities level with an outdoor swimming pool, cinema, gym, and event space on the 22nd floor. The building also has nine levels of below-grade parking, a ground-floor lobby, and two mechanical levels.

The Challenge: Excessive Movement Thornton Tomasetti (TT) and Stark & Ortiz (S+O) were not the original engineers on the project. The tower’s architectural design was substantially complete – and 30 percent of the residential units were already sold – when wind-tunnel testing identified excessive lateral displacements and acceleration in the structure. After an independent peer review, owner Internacional de Inversiones engaged TT and S+O to develop an improved lateral system.

Because many of the residential units had already been sold, changes to the architecture of the apartments would require that the owner give money back to purchasers. The structural challenge was to develop a system that eliminated the excessive movement but fit the building’s predetermined form. The situation also required an extremely aggressive schedule for the re-design and construction.

TT and S+O developed a creative approach to fully address tall-building structural requirements for large vertical gravity and horizontal wind and seismic forces, as well as serviceability concerns of lateral wind and seismic load deflection. Also, the approach addressed STRUCTURE magazine

wind-induced horizontal building acceleration and differential verticalshortening of columns and walls caused by time-dependent concrete creep and shrinkage. A combination of strategies achieved the desired performance while conforming to the capabilities of local contractors to avoid construction delays.

Solutions Virtual Outriggers The slender tower – its height-to-width ratio is 8.7:1 and the central structural core ratio is 19.5:1 – required a lateral-forceresisting system that coupled the central core with perimeter columns. Outrigger trusses are the standard solution, but they can have significant impacts on the architectural program and call for complex structural connections and time-dependent load transfers between columns and core-walls. These impacts all add cost, time, and complexity to construction. The structural engineers developed a “virtual outrigger” system that used stiff diaphragm slabs and reinforced concrete belt-walls at the perimeters of the mechanical spaces on floors 21 (at approximately 40% of the building height) and the 62 (two stories below the roof slab). These elements work together to act indirectly as outriggers, transferring part of the overturning moment from the core to perimeter columns. The stiff floor diaphragms cause the belt walls to tilt, following the core’s rotation, creating a force couple within the columns on opposite sides of the building.

The virtual outrigger system reduced the building period by 20 percent, reduced the core-base overturning moment by 25 percent, and reduced the north/south lateral drift by 30 percent and the east/ west lateral drift by approximately 15 percent. The virtual outriggers efficiently increased structural stiffness without significantly affecting architectural or mechanical layouts.

Concrete link, or coupling, beams over door openings allow the north and south core walls to work together. The beams vary in depth from 1 to 2.75 meters (3.3 to 9 ft), have typical span-to-depth ratios ranging from 1.5 to 3, and use mild-steel reinforcement, except in a

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Construction progress. Courtesy of TIF Digital.

Mat foundation placement. Courtesy of Internacional de Inversiones.

few cases of high shear demand loads, where they are reinforced with structural-steel plate members. Shear walls vary in thickness from 1.05 meters (3.44 ft) at the base to 0.45 meters (1.48 ft) at the top. Thicker walls at the belt levels, which range from 0.6 to 1.6 meters (2 to 5.3 ft) thick, resist the larger shear that results from load reversal caused by the virtual outriggers. The engineers specified three different concrete strengths – 70, 60 and 50 megapascals – to maximize the efficiency of the walls. The diaphragm slabs at the top and bottom of the belt walls are 30 centimeters thick (12 inches) at levels 21 and 22, while 40-centimeter (16-inch) slabs at levels 62 and 63 accommodate concentrated shear stresses at large floor openings for stairs and elevators outside of the core that services the levels above the core termination.

building’s total height. The revenue from these extra floors offset the added cost of the heavier framing system. Costs were also reduced by using post-tensioned waffle-slabs, spanning 35 feet, for the below-grade parking levels where structural mass has minimal impact on lateral performance. In Mexico, material is more expensive than labor, and the 30-centimeter-deep (12-inch) waffle slabs use the same amount of concrete and reinforcement as an equivalent flat-slab only 16.5 centimeters (6.5 inches) deep.

Designing for Fast Local Construction

The time required to redesign Torre Koi’s structural systems made constructability a priority. Local contractors had limited experience in high-rise construction, so it was important to specify materials and Floor Framing for Additional Stiffness systems they could build without costly delays or difficulty.

While the virtual outrigger system significantly improved Torre The aggressive schedule did not allow time for a full material Koi’s performance, it was not sufficient to meet horizontal wind test program for high strength concretes, so the structural engiacceleration limits for occupant comneers designed the structure using fort (18 milli-g at the top residential concrete with a strength at or below floor and 25 milli-g at the top office 70 megapascals – the maximum floor under a 10-year wind), and strength local suppliers could readily there was no room for a tuned mass produce with known and predictable damper. The engineers replaced the properties for stiffness, shrinkage, originally planned one-way beam and creep. and slab floor-framing systems with The engineers performed a nonlinheavier, 25-centimeter-thick (10ear, staged construction analysis with inch) post-tensioned flat-slab floor time-dependent material properties framing. The added mass further in MIDAS Gen, using the known reduced the wind accelerations. properties of the 70-megapascal TT and S+O also worked closely with concrete to capture the long-term the architect and owner to adjust effects of concrete creep and shrinkthe vertical distribution of residential age on the load distribution and and office spaces within the tower. deflected shape of the building.

Moving five levels of less-sensiThis advanced analysis allowed the tive office space to the top of the engineers to maximize column and tower kept residential floors below wall efficiency, minimize materials, levels 59, where wind acceleration and eliminate the need to camber was within industry-standard columns to combat future vertical accepted limits. shortening and compensate construction horizontally. This approach Creative Cost Control reduced costs, time, and complexity While the new flat-slab framing was in the construction process. heavier than the original design, During construction, the conreduced floor thicknesses allowed Full structure (left).Typical floor plans (right). Courtesy of crete sub-contractor instrumented for two additional levels within the Thornton Tomasetti. several columns to measure actual STRUCTURE magazine

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vertical shortening. The data was used to validate and calibrate the structural analysis model. Field measurements correlated very well with the predicted concrete-column strains.

A 4-meter-thick (13foot) mat foundation supported by 77 concrete piles, each 1.5-meter (5 feet) in diameter, supports the tower. Twelve hundred trucks and seven concrete pumps placed the mat’s 7,500 Upper belt wall. Courtesy of Thornton Tomasetti. cubic meters (9,800 cubic yards) of concrete in a continuous pour over 26 hours, making it the largest mass concrete placement in any building in Mexico.

Creative Engineering Delivers Success As the engineering team brought in to redesign Torre Koi’s structure after the architectural design was nearly complete, TT and S+O faced several technical and practical challenges. The structural performance was significantly improved with minimal changes to the existing architectural design and met an extremely aggressive schedule while working with contractors unaccustomed to constructing

tall buildings. How was it done? The team embraced innovative approaches to solve difficult problems, employed advanced analytical capabilities, and collaborated closely with the architect, owner, and contractors to develop successful solutions to the project’s challenges. The result? Mexico’s new tallest building will soon host workers and residents who are embracing new modes of urban life.▪ Chris Crilly, P.E., is a Thornton Tomasetti Senior Associate in the firm’s Washington, D.C. office and was project manager for Torre Koi. He can be reached at CCrilly@ThorntonTomasetti.com.

Mark Tamaro, P.E., is a Senior Principal and office director of Thornton Tomasetti’s Washington, D.C. office. He can be reached at MTamaro@ThorntonTomasetti.com.

Roberto Stark, Ph.D., is President of Stark + Ortiz. He can be reached at stark.roberto@gmail.com.

Project Team Owner: Internacional de Inversiones Structural Engineer: Thornton Tomasetti, Inc. and Stark + Ortiz, S.C. Architect: VFO Arquitectos Wind Tunnel Consultant: Rowan Williams Davies & Irwin Inc. Construction Manager: PMP Consultores

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75 Rockefeller Plaza COLUMN TRANSFERS

By John Hinchcliffe, P.E., Joe Mugford, P.E., and Ramon Gilsanz, P.E., S.E., F.SEI

A

Figure 1. 75 Rockefeller Plaza seen looking north from the ice skating rink.

landmarked, 33-story steel moment frame building, 75 Rockefeller Plaza bookends the north end of Rockefeller Center. The building was originally constructed in 1947 for John D. Rockefellerâ&#x20AC;&#x2122;s Standard Oil Company and features the limestone and cast aluminum façade (Figure 1) that is emblematic of Rockefeller Center. A developer leased the building in 2013 and began a top to bottom renovation. Structural work occurred on every floor and included support framing for new elevators and mechanical equipment. A lobby reconfiguration required removal of four existing building columns on the ground floor. Three of these columns supported existing transfer girders at the second floor. This article focuses on the transfer of these columns, the structural solution constructability, and the load transfer procedure.

The columns scheduled for removal supported the full tributary load of the tower above, with a service gravity load ranging from 1,700 kips to 2,700 kips at the second-floor level. The design team considered several options to remove the columns from the double-height lobby before settling upon new transfer girders below the second-floor level.

Each new transfer girder runs parallel to the existing transfer girder and effectively extends its span from 18 to 30 feet. Figure 2 shows a SecondFloor plan view, and Figure 3 shows an elevation of a new transfer.

Schematic design began before the building could be accessed.

The architect, the steel erector, and the general contractor provided early and consistent input to the initial design. To communicate the structural design concept, Gilsanz Murray Steficek, the structural engineer, printed a three-dimensional scaled model of the existing conditions and the new transfer girders. Each component was printed individually and connected with magnets, simulating the erection and preloading process. When access to the building was gained, and existing field conditions could be verified, lines of communication were already in place between the engineer, the general contractor, and the steel fabricator.

New Load Path

Figure 2. Second-floor framing plan highlighting the columns to be removed in red, the new transfer girders in green, and the existing transfers and column above in black.

STRUCTURE magazine

The new transfer girder cross section (Figure 4) was shaped by architectural constraints, preloading concepts, constructability, and existing conditions. The architectural concept for the lobby, which includes a scalloped ceiling, limited both the depth of the new girder and its bottom flange width.

The new transfer girder was kept below the second-floor slab to preserve rentable square footage. At this level, the new transfer girder needed to fit around the existing girder, which is a pair of built up wide flange shapes.

A composite box girder was used to meet these geometric constraints (Figure 4). Composite box girders are not common in buildings. The American Association of State Highway and Transportation Officials (AASHTO) code was referenced in addition to AISC 360, Specifications for Structural Steel Buildings, for analytical checks and dimensional requirements. The new girder webs are slender and transversely stiffened (Figure 5, page 38). The composite girder steel was sized to elastically

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Figure 3. Elevation of the existing framing, new transfer girder, and jacking rig.

Figure 4. Cross section of the new and existing transfer girders.

resist the dead load, the reduced live load, and the induced load from the column jacking. The plastic capacity of the composite section was sized for the unreduced live load, providing reserve capacity.

The new transfer girder supports the existing structure at the end of the existing transfer girder. Connecting to the existing structure at this location, rather than at the location of the column above, limits the bending moment in the new girder by keeping the point load nearer to the end support. The magnitude of the point load at this location is equal to the end shear of the existing transfer girder, as opposed to the full axial load of the column above. A new built-up channel (Figure 3) picks up the webs of the existing transfer girder and spans between the new transfer girder webs. The connection between the channel and the new girder web was the final connection to be completed after the girder was preloaded and was made with a full penetration weld.

Following the new load path, the existing columns supporting the new transfer girders received approximately 50% more load and required reinforcing. Several columns were directly adjacent to STRUCTURE magazine

existing elevator shafts, and access to reinforce their far flanges would have required shutting down the elevators.

The columns were reinforced on one side with a single cover plate (Figures 6 and 7) to address this constructability challenge. The new cover plate was designed considering the elastic stress in the asymmetric column under a staged construction.

The new box girder webs are 68 inches apart. To connect the new girder to the reinforced column, the column cover plate widens at the top and forks around the existing transfer girder (Figure 7, page 38), where it is stiffened from behind with WTs (Figures 7 and 8, page 38). The box girder connects to the new cover plate with double angle connections, bolted to the girder webs and welded to the cover plate. One design concern at the new girder-column connection was the transfer of moment into the columns when the new girder was loaded and under varying live loads. Strength and Ductility Requirements for Simple Shear Connections under Shear and Axial Load (Thornton, 1997) was referenced to estimate the maximum moment the double angle shear connections transferred to the column. The depth of the new connections was limited to minimize the transferred moment; the reinforced columns and forked plate were designed for this moment.

The existing footings are steel grillages on rock. A geotechnical engineer was engaged to review the increased bearing pressure and it was determined that these footings have the capacity for the additional eccentric load.

The buildingâ&#x20AC;&#x2122;s lateral system is a partially restrained steel moment frame designed under the 1938 New York City Building Code. The 1938 code design wind pressures are much less than current standards. The 1938 Code specified no wind loading for the first 100 feet above grade and a 20 psf load above 100 feet. The base shear at 75 Rockefeller under the 1938 Code is 43% of that under ASCE 7-05 Exposure B. Removing four columns from the lobby and stiffening the adjacent columns modified the buildingâ&#x20AC;&#x2122;s lateral system slightly. Three-dimensional finite element models of the existing and modified moment frames were used to verify the modifications. These models determined the modifications did not significantly change the overall lateral stiffness and did not increase the relative demand on any existing connections by more than 10%.

Load Transfer Pre-loading the new transfer girders using hydraulic jacks provided some advantages. It limited any permanent deflection of the floors above. In a moment frame structure of 75 Rockefeller, preloading

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Figure 5. Half of the box being erected.

Figure 6. Column cover plate erection.

limited the stresses induced in the frame when a single column was allowed to vertically displace. It also allowed the load transfer to be completed while the entire existing structure was in place, maintaining redundancy throughout the loading procedure.

The loading procedure is often dismissed by engineers of record as means and methods of construction handled by the contractor. However, the preloading can have significant structural effects.

Coordinating a procedure with the contractor during the early design stages allowed the engineer to more accurately account for effects like moment transfer at the connections, column strain, and resistance from the existing moment frame.

At 75 Rockefeller, the new girders were jacked at the existing columns to be removed. The rig shown in Figure 3 and Figure 8 was used to push the existing column up and pull the new girder down.

Rivets were removed from the existing splice plates above the first floor in the column to be removed (Figure 7 ). To transfer the load, the axial strain in the existing columns needed to be assessed. The existing columns elongated as they were unloaded (approximately 1⁄8 inch); similarly, the reinforced columns shortened as they received additional loads (approximately 1⁄16 inch each). To achieve a separation at the splice, the columns being removed must displace by the sum of these two values relative to the columns receiving the load. This displacement was resisted by 30+ stories of the existing moment frame above. Achieving these displacements induced a stress (approximately equivalent to a 200 kip point load) into the moment frame and the new girder.

Incrementally loading the girder, and maintaining a comprehensive record of displacements and jack pressures, helped to confirm design assumptions and was invaluable for informing decisions made in the field. Each girder was loaded to the full jacking load

Figure 7. Transfer girder erection.

without a visually observable separation at the column splice.

Displacement records helped inform the decision on when to make the final connection.

Monitoring was continued when the jacks were released and while the column was cut. This proved important, as the channel of the final connection (Figure 3) deflected when it received load, relieving some pre-stress from the girder and resulting in an upward girder displacement of approximately 1⁄16 inch. Because of this, the column was cut with some load in it. The web of the column was removed first. The built-up flanges were then cut in a cradle starting at the toes and cutting incrementally to create a seat. The cut at one toe was started higher than the other so that they did not meet in the middle but overlapped, letting the flange of the column yield in bending as the final displacement (of just over 1⁄16 inch) occurred.

Conclusions Altering an existing structure load path is challenging, and the manner in which it is done can significantly affect the structure. As engineers, we can produce better designs by coordinating the load transfer procedure early in the design, analyzing the effects at each stage of the procedure and carefully monitoring the behavior of the existing structure.▪ The online version of this article contains references.

Please visit www.STRUCTUREmag.org.

John Hinchcliffe, P.E., is a Senior Engineer at Gilsanz Murray Steficek LLP Engineers and Architects. He may be reached at john.hinchcliffe@gmsllp.com.

Joe Mugford, P.E., is an Associate at Gilsanz Murray Steficek LLP Engineers and Architects. He may be reached at joe.mugford@gmsllp.com.

Ramon Gilsanz, P.E., S.E., F.SEI, is a Founding Partner at Gilsanz Murray Steficek LLP Engineers and Architects. He may be reached at ramon.gilsanz@gmsllp.com.

Project Team Owner and General Contractor: RXR Realty Structural Engineer: Gilsanz Murray Steficek Architect: Kohn Penderson Fox Steel Erector: Orange County Ironworks Figure 8. Jacking rig.

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Structural DeSign design issues for structural engineers

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oday’s design, engineering, and construction professionals face a host of complex construction challenges.

Industry initiatives focus on building a safer, more sustainable built environment. Owners seek maximum return on investments. Engineers pursue innovative technologies that maximize asset life and minimize life-cycle costs. Construction teams focus on quality, efficiencies, and costeffectiveness. Aligning priorities and achieving design intent is a delicate balance of cooperation, collaboration, and communication.

A vast majority of construction projects incorporate the use of concrete for its versatility, durability, and sustainability. It has become the single most widely used building material in the world. With costs estimated at over $4 Trillion to repair or reconstruct America’s deteriorating built environment, ensuring this essential component provides the quality and intended performance for its designed service life is crucial.

A key challenge in concrete design is to address cracking – the primary reason deterioration and corrosion occur in concrete. Ultimately, cracks in the concrete allow moisture, salts, and other contaminants to reach the metal surface of reinforcement and cause corrosion, deterioration, and failure. Permeability is another critical factor that provides entry points and pathways for contaminants to penetrate into the concrete and begin their destructive work.

Though there isn’t a “one size fits all” solution, Type K shrinkage-compensating cement is a proven method of improving the concrete element at its core – by improving the cement paste itself. It was specifically designed to help overcome key challenges commonly experienced with concrete installations and has been successfully used in concrete design since the early 1960s. It has a notable history of success in post-tensioned structures, chemically pre-stressed concrete, slabs-on-grade, concrete containment, and cast-in-place elements where higher performance and extended service life were required (Figure 1).

Shrinkage-Compensating Concrete Designs Designing Best Value Solutions

By Susan Foster-Goodman and Ken Vallens

Susan Foster-Goodman is Director – Strategic Initiatives & Komponent Sales at CTS Cement Manufacturing Corp. She can be reached at sgoodman@ctscement.com.

Ken Vallens is President and CEO of CTS Cement Manufacturing Corp. headquartered in Cypress, CA. He can be reached at kvallens@ctscement.com.

(CSA) cement technology, which has several distinguishing characteristics differentiating its performance: • It contains no C3A, so sulfate resistance is maximized. When combined with Type V portland cement, maximum sulfate resistance can be achieved in shrinkage-compensating concrete designs.

• It has a high C2S content, which means longterm strengths can be maximized.

• Its early strength element, CSA, has a unique hydration mechanism that ensures efficient consumption of mix water and improves the overall quality and performance of the concrete.

During the hydration process, water molecules are chemically retained (“bound”) within a tightly woven network of ettringite crystals, effectively eliminating excess bleed water. This hydration mechanism affords many performance advantages: - By preventing the egress of excess mix water, voids and capillary channels typically left behind are prevented. These voids would have created space for drying shrinkage, volume change, and easy entry points for contaminants. By preventing these voids, drying shrinkage challenges are overcome, dimensional stability is improved, and lower permeability is achieved (Figure 2).

- Elimination of bleed water helps maintain the integrity of the water/cement ratio at the surface of the concrete, which improves abrasion and impact resistance and prevents laitance and other debris in the concrete from being drawn to the surface.

- The ettringite crystals formed during hydration create expansion within the concrete.

The proper dosage of the additive ensures the designed expansion is achieved to overcome the shrinkage characteristics of the local portland cement and aggregates.

- This advanced hydration mechanism effectively overcomes hydral volume change, prevents drying shrinkage cracking, overcomes

What It Is Type K cement is a blended cement that incorporates an expansive cementitious additive chemically engineered to overcome drying shrinkage cracking, reduce permeability, and improve sulfate resistance. This additive (marketed as Komponent®) is Figure 1. Proven Type K Shrinkage-Compensating Cement powered by calcium sulfoaluminate performance for over 50 years.

40 January 2018

any concrete structure without changing the reinforcing or joint detailing.

Designers should familiarize themselves with ASTM C845, Standard Specification for Expansive Hydraulic Cement, and ASTM C806, Standard Test Method for Restrained Expansion of Expansive Cement Mortar, and review the principles, methods, and details related to structural design with shrinkage compensating concrete provided in ACI 223.

Shrinkage specifications are tightening (with a maximum 0.02 or less specified more often), and associated performance time frames are being extended to ensure that designed service life expectations are achieved. Type K cement was engineered to meet and exceed these rigorous requirements and offers a proven solution for today’s engineering community.

Figure 2. Egress of excess bleed water leaves behind room for drying shrinkage.

restraint-to-shortening challenges, reduces the loss of pre-stress due to concrete shrinkage in post-tensioned designs, lowers permeability, and ensures the long-term dimensional stability of the concrete.

Type K cement provides a higher quality cement paste that can play a significant role in maximizing durability, performance, and service life of concrete designs.

How It Works Shrinkage-compensating cement is engineered to work in conjunction with internal restraint.

The expansion, which is controlled by the reinforcement, puts the reinforcement in tension (positive steel strain) and the concrete in compression early in the process. The expansive forces developed during the 7-day wet cure period create controlled compressive stresses that keep concrete in compression throughout its service life. It is similar to placing a concrete bar in a powerful clamp, in which the length of the bar fits tightly within the perimeter limits of the clamp. When the bar is heated, it is unable to expand due to the restraint of the clamp.

The bar is now in compression.

Drying shrinkage in concrete is similar to cooling the concrete bar. As the bar cools, the compression is reduced until the bar reaches its original temperature and length.

Further cooling of the concrete bar would result in additional shortening, creating a gap between the ends of the bar and the clamp.

Commonly in concrete, shortening results in drying shrinkage cracking (Figure 3).

With shrinkage-compensating cement, expansion maximizes restraint forces and offsets (or “compensates for”) detrimental tensile

Figure 3. How it works.

stresses and negative strains caused by drying shrinkage that result in shrinkage cracking and, ultimately, contamination, corrosion, early deterioration, and failure.

Industry Standards and Guidelines The graph provided in ACI 223, Guide for the Use of Shrinkage-Compensating Concrete (Figure 4, page 42), provides a good representation of what is happening. During the 7-day wet cure period, the expansion effects of the ettringite formation are evident as water molecules are consumed. This puts the concrete into compression and the reinforcement in tension early. The goal is to dose the mix design sufficiently to overcome the anticipated drying shrinkage of the portland and aggregates and achieve a net zero shrinkage.

ASTM C878, Standard Test Method for Restrained Expansion of Shrinkage-Compensating Concrete, assists engineers in determining the most effective dosage for the project. Typically, a 15-17% replacement of cement content is sufficient to achieve shrinkage compensation. A minimum ratio of reinforcement to the gross concrete area is 0.0015, with no maximum specified. Though the “[c]oncrete member expansion is reduced as the amount of reinforcement is increased, the amount of compressive stress in the concrete will be increased” (ACI 223, Figure 2.5.3).

Expansion results must achieve a 7-day expansion between 0.04% to 0.1%, with expansion at 28 days not to exceed 0.15% of the 7-day expansion. These limits were chosen to ensure the ability to substitute shrinkage compensating cement within this range for

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Project Advantages When evaluating construction dynamics, project efficiencies, and cost impacts, there is a host of advantages that can be realized when using shrinkage-compensating concrete designs. Dimensional stability is maintained when using Type K. This means costly, time-consuming solutions for restraint-toshortening challenges (e.g., pour/delay strips, expansion joints, slip joints, additional reinforcement) can be eliminated.

Type K shrinkage-compensating concrete achieves 30-40% greater abrasion resistance due to its advanced hydration mechanism.

Slab curling and unsupported slab edges that result in cracked corners and spalls are prevented. The time and expense of slab jacking, grinding, and leveling prior to finish installation can be avoided, and designed Floor Flatness (FF) is maintained.

Larger, more monolithic slab and wall placements can be achieved, with extended joint spacing and length-to-width ratios up to 3:1 are achievable. Containment wall panels can be successfully placed at up to 130 feet.

Extended bridge deck joint spacing can be achieved up to 300 feet. Control joint spacing up to 150 feet in slab designs can be accomplished with up to a 95% reduction in joints to place, cut, treat, and maintain.

Fewer joints improve efficiencies of seamless flooring installation, with few if any joints to bridge. It also contributes to satisfying new OSHA requirements for worker safety by minimizing respirable crystalline silica exposure due to saw cutting concrete.

Larger placement sizes also reduce load transfer reinforcement and minimize mobilizations and formwork. In concrete containment designs, more monolithic placements minimize leakage and seepage points, reduce water

Figure 4. Initial expansion compensates

Figure 5. Understanding Type K Cement vs. SRAs.

stops and structural joint detailing, ease construction in “no joint zones,” and eliminate checkerboard casting. Type K’s initial high slump makes it ideal for high volume pump productivity and produces a smooth, dense surface for sanitary engineered structures.

By compensating for the effects of drying shrinkage through expansive cement technology, temperature and shrinkage steel can be reduced or eliminated, making thinner slabs and walls viable.

Much can be achieved by addressing concrete challenges at their core. By improving the cement paste itself, Type K cement provides unmatched versatility in design and application, and a wide range of opportunities to improve project efficiencies and realize cost savings.

Installation Considerations Type K cement is compatible with traditional complementary products like plasticizers, air entrainment, corrosion inhibitors, retarders, water reducers, and hydration stabilizers. As with traditional portland cement concrete, best practices must be used for hot and cold weather concreting. Common means and methods for placement and finishing are used for Type K cement concrete. Due to the lack of excess bleed water, finishing can begin sooner.

Commercial (4000+ psi), industrial (6000+ psi), and heavy-duty (10,000+psi) strength requirements can be achieved. Supplementary cementitious materials (SCM) (e.g., fly ash, silica fume, slag cement) can be used up to a maximum 18% by weight of cement. Most integral waterproofing additives are compatible with Type K cement, though they should be tested to ensure designed expansion is not affected.

As a result of the consumption of water molecules used to create ettringite crystals during hydration, water demand can be slightly higher with Type K cement, though compressive strengths and flexural strengths are comparable to portland designs. Low w/c ratios can be achieved with careful consideration of various other mix design factors. Consulting with the Type K cement manufacturer for low w/c ratio recommendations is advised.

Water-curing is essential with shrinkagecompensating concrete designs to ensure the designed expansion is achieved. Common methods include polyethylene sheets, absorptive moisture-retaining covers, ponding, or the use of soaker hoses or sprinklers with a curing blanket.

Production can be achieved at the batch plant or on-site via portable silos. Slurry machines can also be used when production facilities or project dynamics necessitate versatility in introducing the Type K additive or based on contractor preference.

Cost Considerations

Figure 6. Key construction advantages of Type K Cement.

When evaluating project costs, a thorough review of project scope and the various opportunities available for savings in design, detailing, time, labor, and materials during construction should be performed. An evaluation of lifecycle costs, operational efficiency expectations, and return on investment requirements of the

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owner must be considered. When performance requirements necessitate augmenting designs to ensure durability, structural integrity, minimal maintenance, or extended asset life, preventative actions must be incorporated into the design and project budget. Type K cement provides a high-performance, cost competitive, and value-added solution that affords a host of additional project savings potential.

Type K cement is a blended cement consisting of the Type K additive, local portland cement, and aggregates. Though initially only available as a pre-blended cement, industry standards, specifications, test methods, and guidelines have been developed and proven, making the use of a concentrated additive viable. The concentrated additive is the most common and economical method used today to create Type K cement in even the most remote locations. Combining the Type K additive with local portland cement and aggregates provides a cost-effective way to produce Type K shrinkage compensating concrete and grout.

It affords flexibility in production methods to suit unique project conditions. The Type K additive is available in bulk rail, bulk truck, bulk bag, and small bag units, and is offered via a direct-to-market sales channel.

Conclusion With more restrictive shrinkage specifications on the rise and a growing emphasis on maximizing design life and minimizing lifecycle costs, Type K shrinkage-compensating concrete is enjoying a resurgence in popularity. It offers a “Best Value” solution that supports professionals throughout the industry in achieving primary objectives of building a safer, more sustainable and economical built environment.▪ The online version of this article contains detailed references. Please visit www.STRUCTUREmag.org.

Anchor Updates DEWALT

Phone: 845-230-7533 Email: mark.ziegler@sbdinc.com Web: www.dewalt.com Product: AC200+ Adhesive Anchoring System and Post-Installed Rebar System in Concrete Description: The AC200+ is a two-component, high strength adhesive anchoring and post-installed rebar system. The AC200+ is approved for bonding threaded rod and reinforcing bar hardware into drilled holes in cracked and uncracked concrete base materials. Evaluated and recognized for sustained loads, freeze-thaw performance and seismic connections. ICC-ES ESR-4027.

Dlubal Software, Inc.

Headed Reinforcement Corporation (HRC)

Phone: 714-852-1333 Email: Jeremy@hrc-usa.com Web: www.hrc-usa.com Product: HRC 670 HeadLock™ Description: The HRC HeadLock is a field installed head that allows you to anchor rebar when the standard hook has been installed at the wrong elevation. Cut off hook to correct elevation, press on HeadLock, and snap off single bolt.

Simpson Strong-Tie

Phone: 800-925-5099 Email: web@strongtie.com Web: www.strongtie.com Product: Stainless-Steel Titen HD® Heavy-Duty Screw Anchor Description: The new stainless-steel Titen HD screw anchor can now be installed in exterior and corrosive environments. It is made of Type 316 stainless steel with serrated carbon-steel threads at the tip. The innovative design effectively cuts the concrete while reducing the carbon steel in the anchor to maximize corrosion resistance.

Product: SET-3G™ High-Strength Anchoring Adhesive Description: SET-3G is the latest innovation in epoxy adhesives. The high-strength anchoring adhesive can be installed in extreme concrete temperatures from 40°F to 100°F, as well as in dry or water-filled holes in concrete. SET-3G provides the high bond strength values needed for a variety of adhesive anchoring applications.

Trimble

Phone: 770-426-5105 Email: jodi.hendrixson@trimble.com Web: www.tekla.com Product: Tekla Structural Designer Description: Revolutionary software that gives engineers the power to analyze and design steel and concrete buildings efficiently and profitably. Physical, information-rich models contain all the intelligence needed to fully automate the design and document your project, including all end force reactions communicated with two-way BIM integration, comprehensive reports and drawings.

Product: Tekla Structures Description: Tekla is an Open BIM modeling software that can model all types of anchors required to create a 100% constructible 3D model. Anchors can either be created inside the software or imported directly from vendors that have 3D CAD files of their products.

MacLean Power Systems

Phone: 731-330-4025 Email: jtully@macleanpower.com Web: www.MacleanDixie.com Product: Helical Piles and Driven Anchors Description: MPS manufactures the highest quality helical piles, helical tiebacks and soil nails along with driven anchors. Our Duckbill and Manta Ray anchors have set the industry standard for driven anchors, while engineers and contractors alike demand MacLean helical piles and anchors in their specifications.

All archive articles available online: www.STRUCTUREmag.org.

RISA

Phone: 949-951-5815 Email: info@risa.com Web: risa.com Product: RISAConnection Description: Get baseplate and anchor design for RISA-3D and RISAFloor models using RISAConnection. Column base reactions, including biaxial moments, are sent for each load combination to RISAConnection. This is more accurate than traditional methods of designing for envelope reactions. With RISA-3D and RISAFloor integration, RISAConnection is the new standard for baseplate design.

Listings are provided as a courtesy. STRUCTURE® magazine is not responsible for errors.

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January 2018

TILT-UP CONSTRUCTION SOLUTIONS

Dayton Superior has been synonymous with quality, safety, professional service, and an unmatched tilt-up product portfolio.

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TILT-WERKS.COM

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Phone: 267-702-2815 Email: info-us@dlubal.com Web: www.dlubal.com Product: RFEM Description: Effectively analyze complex connection layouts utilizing surface and solid elements, automatic finite-element meshing, mesh refinements, and surface intersection capabilities. Perform the required ultimate and serviceability limit state designs of reinforced concrete according to ACI and other international standards. Accurately represent soilstructure interaction with multiple soil layers to obtain calculated foundation stresses and settlements.

news and information from anchor companies

InSIghtS new trends, new techniques and current industry issues

A

key aspect of reinforced concrete design is the calculation of the required reinforcing area and the required development and splice lengths for proper reinforcement details. In most structures, the reinforcement is cast with the concrete and subsequent concrete pours, i.e. “new” concrete, is placed around the cast-in-place reinforcement protruding from the existing concrete.

The cast-in-place reinforcement’s development length is prescribed by ACI 318-14 Chapter 25, “Reinforcement Details,” or Section 25.4.2.3, (Chapter 12, Section 12.2.3, in previous versions of ACI). Engineers are familiar with the basic expression for tension development length in Ch. 25 from the following equation:

Post-Installed Adhesive Anchor Systems Reinforcing Bar Development Length using the Provisions of ACI 318

By Christopher Gamache, P.E.

Unfortunately, reinforcement is not always cast into the existing concrete when a new section of concrete will be connected to existing concrete.

Examples include renovations where the original building plan did not account for the new concrete section, strengthening a structure to improve the seismic performance, or rehabilitation of an existing structure.

When cast-in reinforcement is not provided in the existing concrete, a post-installed adhesive anchor system is commonly used to attach reinforcement to the existing concrete. Post-installed adhesive anchor systems have been around since the mid-1970s and are widely used in construction today for a variety of applications including attachment of threaded rods and reinforcement to concrete. Design of post-installed adhesive systems is regulated by ACI 318-14 Chapter 17, “Anchoring to Concrete.” (Note, the provisions of ACI 318-14, Chapter 17 are the same as the provisions of ACI 318-11, Appendix D.) However, Chapter 17 is intended for the design of both cast-in-place headed bolts and post-installed anchors to transfer tension and shear loads from a connected structural element, typically a steel fixture, to the concrete member and does not deal with reinforcement development lengths.

Additionally, Chapter 17, Section 17.3.2.3 only permits designs for a maximum embedment of 20 times the anchor diameter, which can be less than the embedment depth calculated for development.

So how would an engineer design an application where attaching new concrete to an existing concrete structure is required when there are no cast-in-place bars to connect having proper development length?

To address this, ICC Evaluation Services updated the test criteria, Acceptance Criteria for Post-Installed Adhesive Anchors in Concrete Elements (AC308), in June 2013. The update includes adhesive anchor system qualification tests for use with the development length provisions of Chapter 25 and seismic requirements of ACI 318-14 Chapter 18, “Earthquake Resistant

Chris Gamache is the Manager of Approvals and Project Engineering/ Anchors for Hilti North America.

He is responsible for creating the technical data for the Hilti North American Product Technical Guide, Volume 2, for Anchor Fastening, and publishing external evaluation reports such as ICC-ES ESRs. He can be reached at christopher.gamache@hilti.com.

ICC-ES AC308 Figure 4.6 – test specimen to confirm bond and splitting behavior at deep embedment. Courtesy of 2016 ICC Evaluation Service, LLC.

44 January 2018

Structures.” Before 2013, AC308 only included tests for anchor designs applicable for ACI 318-11 Appendix D. In 2013, AC308 Table 3.8, “Test program for evaluating deformed reinforcing bars for use in post-installed reinforcing bar connections,” was added for manufacturers that wanted their adhesive system to be recognized for use with the reinforcement development provisions. Successful completion of the tests show that the adhesive anchor system will perform equally to a cast-in-place reinforcing bar, even with an embedment depth beyond the 20 anchor diameters maximum in Chapter 17. The testing is rigorous and is intended to prove the following: • The adhesive anchor system has a stiffness that will transfer load over the full length of the reinforcement and a bond strength that is equal to the cast-in reinforcement tension strength.

• The adhesive anchor system can be installed at embedment depths beyond 20 times the bar diameter without the adhesive product hardening before the reinforcement is inserted, and the adhesive injection is free of voids.

To show equal performance and stiffness, the bond/splitting force test found in AC308

Table 3.8 requires two reinforcing bars to be cast in opposite corners of a concrete “column” and two post-installed adhesive anchors with reinforcing bars to be installed in the other corners. The embedment depth is 35 bar diameters. Each cast-in and postinstalled reinforcing bar is tension tested independently. The mean ultimate tension load and displacement must be equivalent between the two systems. The test setup for the bond/splitting force test is shown in Figure 4.6 of AC308 (included in this article). This test checks the load transfer of the adhesive system over the entire length of the reinforcing bar.

To show that the adhesive anchor system can be installed properly at embedment depths beyond 20 bar diameters, AC308 Table 3.8 has two installation verification tests at an embedment depth of 60 bar diameters.

The first test verifies that the reinforcing bar can be inserted to the full 60 bar diameter embedment before the adhesive system begins to gel (harden). All adhesive systems have product-specific working (gel) times and cure times. Adhesive anchors that can be installed at the 20 bar diameter embedment for anchor applications may not work at this extended length if the product has a

fast gel time. The second test verifies that the adhesive system, when injected, leaves no voids in the drilled hole which would reduce the bond strength.

After successful completion of the AC308 Table 3.8 tests, the product can be recognized in a product-specific evaluation report from an independent product evaluation service such as ICC Evaluation Services.

Evaluation reports should specifically mention testing is in accordance with AC308 Table 3.8. Engineers can now use Equation 25.4.2.3a from ACI 318 to design reinforcing development lengths for reinforcing bars whether cast-in or installed with a qualified post-installed adhesive system.

In summary, where cast-in reinforcing bars are not present in the existing concrete member, structural engineers can use the traditional ACI 318 reinforcing bar development lengths when a qualified post-installed adhesive system is used. The development length design will be the same regardless if the reinforcement is cast-in or post-installed.

The selection of the proper adhesive can be authenticated with an independent evaluation report that verifies the product has successfully passed the difficult AC308 Table 3.8 test requirements.▪

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Build your NCEES Record today. ncees.org/records

January 2018

Structural SuStainability

sustainability and preservation as they pertain to structural engineering

Structural Engineers and Climate Change

By Megan Stringer, S.E., LEED AP BD+C and Mark D. Webster, P.E., LEED AP BD+C

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nthropogenic (human-caused) emissions of greenhouse gases (GHG) into the atmosphere, particularly carbon dioxide, are causing dramatic changes to the earth’s climate and oceans. These changes are described in detail in the authoritative and comprehensive reports written by the United Nations Intergovernmental Panel on Climate Change. Through building construction and building use, the construction industry contributes nearly 40% of total U.S. anthropogenic carbon emissions to the environment. As new building operational systems become increasingly energy efficient, the relative contribution of embodied carbon impacts to the overall carbon equation, including structural framing, becomes increasingly important. A new technical report by the ASCE SEI Sustainability Committee, Structural Materials and Global Climate: A Primer on Carbon Emissions for Structural Engineers, noted how structural engineers could play a key role in reducing total construction industry emissions through informed design and efficient structures. The following article summarizes this report.

Climate Change Basics

Contribution of Buildings to Climate Change Environmental impacts from buildings are separated into two categories; operational impacts that occur during service life and embodied impacts that occur during construction, renovation, and demolition.

Although embodied impacts are often assumed to account for 20% or less of total impacts and building operations for the remaining 80% plus, these percentage breakdowns may be misleading. First, they are based on a building lifespan of 50 years or greater, and many buildings are either refurbished or torn down before reaching that age.

Second, buildings are becoming more energy efficient and striving for net zero energy use.

Embodied impacts are therefore becoming increasingly significant.

It is imperative that we not only reduce emissions from building operations but reduce embodied impacts as well. Materialrelated GHG emission reductions will have a near-term impact on climate change, when it is most needed, whereas reductions

due to operating energy improvements are spread out over decades. Structural engineers need to be aware of the embodied carbon impacts of their design decisions and consider design choices to reduce emissions.

By utilizing life cycle assessment, structural engineers can quantify the environmental impacts of the structures they design.

See the ASCE-SEI technical report for more information.

Structural Material Contribution The climate change impacts associated with the primary structural materials are briefly introduced below. Again, the ASCE-SEI technical report contains more information.

Concrete The four basic components common to all concrete are cementitious materials, water, and coarse and fine aggregate. Although the cementitious material makes up the smallest proportion of a concrete mix design, it contributes approximately 90% of its carbon footprint when it is ordinary portland cement.

This high contribution is attributable to both the energy used in the manufacture of portland cement and the chemical reaction, called calcination, associated with the conversion of limestone to cement. Other cementitious materials such as slag cement and fly ash are commonly used and significantly reduce the carbon footprint. The other processes in concrete production (mining, crushing of aggregates, and transportation of aggregates/cement) contribute the remainder of concrete’s carbon footprint.

Temperature Anomaly (C)

Climate change is defined as a change in global or regional climate patterns. Since the mid-to-late 20th century, the earth has been experiencing changes in global climate from the effects of global warming. This is primarily attributed to increased levels of GHG in the earth’s atmosphere, the most abundant of which is carbon dioxide (Figure 1). Since

1900, the global average temperature has risen 1.5 degrees Fahrenheit, with most of the increase occurring in the last 40 years (Figure 2). This relatively sudden rise in temperature is destabilizing the earth’s weather systems, causing new rainfall and drought patterns, more rapidly rising sea levels, more heat waves, increasing ocean acidification, and the potential for higher energy storms.

The same trend can be seen if you look at the carbon dioxide levels in the atmosphere during the last three glacial cycles.

Figure 1. Carbon dioxide in atmosphere. Courtesy of NASA/NOAA. Figure 2. Global land-ocean temperature. Courtesy of NASA/NOAA.

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Steel The majority of structural steel consumed in the U.S. is produced and fabricated domestically. It is produced from one of two methods: the electric arc furnace (EAF), or the basic oxygen furnace (BOF). The majority of steel’s carbon impact comes from the manufacturing process used. The EAF process creates half of the carbon dioxide emissions of the BOF process. One reason for this is the high recycled content (over 90%) of the raw material used by EAFs. This process is used for domestic hot-rolled structural steel and rebar production, while hollow structural sections and cold-formed steel are currently produced using both processes.

report. In addition to the material-specific strategies, broader approaches to reduce GHG emissions can be made, including: • reducing material quantities • mitigating thermal bridging • utilizing structure as finish • performance-based design • using alternate structural systems than typically used for a particular building type • utilizing functionally equivalent materials

• sourcing salvaged materials • designing for deconstruction All of these strategies, both at a material level and building level, must be taken with a holistic view to have the most beneficial impact. A designer needs to consider the environmental tradeoffs of design decisions, as well as the amount of potential saving a strategy offers and the timeframe for those savings.▪

Wood

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Wood “sequesters” carbon; as trees grow, they convert carbon dioxide from the air into wood fiber. As long as that wood fiber remains intact, the carbon is stored in it. When it burns or decays, the carbon is returned to the atmosphere. The carbon impact of wood comes from harvesting, milling, transportation, and the forest management.

The role of forest management practices is difficult to quantify and is the subject of ongoing study.

Reducing Environmental Impacts of Buildings Strategies for reducing the carbon impacts of the individual structural materials are discussed in the ASCE SEI Sustainability Committee technical Megan Stringer is a Senior Engineer at Holmes Structures’ San Francisco Office. She is an ASCE SEI Sustainability Committee Carbon Task Group Member and past chair of the SEAOC Sustainable Design Committee. She can be reached at mstringer@holmesstructures.com.

Mark D. Webster is a Structural Engineer at Simpson Gumpertz & Heger Inc.’s Boston-area office and is chair of the ASCE SEI Sustainability Committee Carbon Task Group. He can be reached at mdwebster@sgh.com.

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Codes and standards updates and discussions related to codes and standards

C

hanges to the 2018 International Building Code (IBC) and 2018 International Existing Buildings Code (IEBC) were approved by the International Code Council (ICC) during their 2015/2016 code development cycle. This article outlines changes to the code requirements for wood construction, the majority of which are changes to the IBC. Only a few changes for exterior balconies involve the IEBC. For this article, all changes noted are to the IBC; any changes to the IEBC will be specifically called out. Accompanying the discussion of each code change is the ICC code change tracking number [colored/bracketed] that can be used to search for more information on the ICC website (iccsafe.org). Appendix A, available in the online version of this article and at www.awc.org, contains a strikethrough/underline format of changes where it is deemed helpful to understand the code changes mentioned herein.

2018 IBC and 2018 IEBC Changes Related to Wood Construction John “Buddy” Showalter, P.E., David P. Tyree, P.E., C.B.O., and Sandra Hyde, P.E.

David P. Tyree is the Central Regional Manager and John “Buddy” Showalter is Vice President of Technology Transfer for the American Wood Council (AWC). Sandra Hyde is Senior Staff Engineer with the International Code Council. Contact Mr. Showalter (bshowalter@awc.org ) with questions.

Referenced Standards American Wood Council (AWC) standards, as well as other code referenced standards, are updated [ADM94-16]. The 2018 National Design Specification® (NDS®) for Wood Construction and the 2018 Wood Frame Construction Manual for One-and-Two Family Dwellings were “Approved as Submitted” without modification.

The 2015 Special Design Provisions for Wind and Seismic (SDPWS) and 2015 Permanent Wood Foundation Design Specification are both still referenced in 2018 IBC. The following updated APA-The Engineered Wood Association ANSI standards were also included [ADM94-16]: • ANSI A190.1-2017 Structural Glued Laminated Timber • ANSI/APA PRG320-2017 Standard for Performance-Rated Cross-Laminated Timber • ANSI/APA PRP 210-2014 Standard for Performance-Rated Engineered Wood Siding • ANSI/APA PRR 410-2016 Standard for Performance-Rated Engineered Wood Rim Boards

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• ANSI 117-2015 Standard Specification for Structured Glued Laminated Timber of Softwood Species

Approved Agencies Certification report writing agencies were introduced into the definition of Approved Agency in Section 202. [ADM6-16, Part 1 AMPC1]

Exterior Balconies • Clarifies removal of balconies from the scope of IBC Chapter 14 Exterior Walls since all balcony provisions were moved to IBC Chapter 7 Fire and Smoke Protection Features, IBC 705.2 and 706.5.2 [S1-16 AM] • Requires that ventilation openings be provided similar to rafter spaces when the floor structure of exterior balconies and decks are enclosed [S7-16 AM] • Incorporates the requirement from ASCE 7 for design live load of balconies and decks at 1.5 times the live load of the area served by the balcony or deck – not required to exceed 100 psf [S85-16] • Requires detailing on plans of all impervious moisture barrier system elements (including manufacturer’s instructions when applicable) if the impervious moisture barrier option is used in IBC 2304.12.2.5 for wood framing supporting weather-exposed permeable floors, such as concrete or masonry slabs [ADM 77-16] • Requires inspection of all impervious moisture barrier system elements, or special inspection can be utilized at the option of the code official if the impervious moisture barrier option is used in IBC 2304.12.2.5 for wood framing supporting weather-exposed permeable floors [ADM 87-16] • Requires the impervious moisture barrier system to have positive drainage of water that infiltrates the permeable floor above the impervious moisture barrier when that option is used in accordance with IBC 2304.12.2.5 [S279-16 AMPC1]

• Clarifies the IBC 2304.12.2.2 treated wood exception for posts supported on pedestals [S278-16 AM] • Modifies IBC equations 23-1 and 23-2 for deflection of diaphragms and shear walls fastened by staples to be consistent with AWC SDPWS equations for nailed diaphragms and shear walls [S282-16 and S284-16] • Clarifies in IBC 2308.2.3 that buildings with slab-on-grade floors can exceed a floor live load of 40 psf and still use the conventional wood frame construction provisions of IBC 2308 [S287-16] • IEBC Chapter 1 revisions require details in construction documents and inspections for impervious moisture barriers used in exterior balconies [ADM 77-16 AMPC1].

An article on balcony detailing is available at this link: http://bit.ly/2AtN5k6

Special Inspection and Structural Observation • For structural observation, modifies the wind trigger from 110 mph to 130 mph for Risk Categories III or IV to match the current factored level of wind forces [S133-16] • Clarifies the main wind force-resisting system fastening exception to special inspection in wood frame construction (based on nail spacing for sheathing exceeding 4 inches on center) at the panel edges. [S145-16 AM]

Other Changes • Clarifies the definition of Light-frame Construction by removing “method of construction” from the definition [G2-16 AM] • Revises Table 1604.3 Deflection Limits in footnote “d” to recognize different wood products’ creep behavior; specifically – seasoned lumber, structural glued laminated timber, prefabricated wood I-joists, SCL, cross-laminated timber and wood structural panels [S63-16 AM and S67-16]; includes correlating change to add roof live load to the load combination • Clarifies that hardboard siding used structurally must conform to ANSI A135.6 and be identified by a label containing the approval agency [S258-16] • Creates consistency with International Residential Code (IRC) wood structural panel roof sheathing nail size by inclusion of 8d common nail and adds the Roof Sheathing Ring Shank Nail (RSRS-01) to Table 2304.10.1 as options for roof sheathing attachment [S272-16]

• Corrects the 10d common nail length, removes redundant requirements for stud nailing, creates consistency with the IRC for roof sheathing attachment, and adds an option for deformed shank nail roof sheathing attachment [S273-16] • Adds a reference to IBC 2304.9 for lumber decking in IBC 2304.11 for heavy timber [S276-16] • Adds an alternative fastening schedule for the construction of mechanically laminated decking made from 2-inch nominal dimension lumber to IBC 2304.9.3.2 [S281-16] • Corrects the staple description for stapled fiberboard shear walls in Table 2306.3(2) [S286-16] • Updates Table 2308.4.1.1(2) for Southern Pine No. 2 in lieu of Southern Pine No. 1 for interior bearing wall girder and header spans, and includes the dropped and raised header distinction for spans [S288-16] • Updates Table 2308.4.1.1(1) with Southern Pine No. 2 in lieu of Southern Pine No. 1 for exterior bearing wall girder and header spans and includes the dropped and raised header distinction for spans [S289-16] • Adds prescriptive framing and connection requirements to IBC 2308.5.5.1 for single member (single ply) headers, consistent with the IRC, and coordinates code charging language with existing connection tables [S292-16] • Updates references to current AWPA section numbering for preservative treatment used in permanent wood foundations and for wood shakes [S40-16] • Clarifies that the minimum 5 psf horizontal live load applies to firewalls [S55-16 AM] • Clarifies IBC 1615.1 regarding the applicability of structural integrity provisions in high rise buildings and precludes misinterpretation regarding frame buildings [S126-16] • Revises IBC 1810.4.1.5 to require the removal of timber piles when a substantial and sudden change in rate of pile penetration occurs during driving [S233-16]

Fire Retardant Treated Wood (FRTW) • The approved modifications to IBC 2303.2.2 clarify that the “other means during manufacture” subsection is not intended to permit surface-protected products as outright replacements for fire retardant treated wood (FRTW), given the requirement for chemical impregnation into the wood. The modifications also preclude interpreting IBC 2303.2.2 as a ban or prohibition on surface-coated products.

As has been the case for some time, wood products protected by surface treatments can be evaluated and approved by using the provisions of IBC 104.11 [S262-16 AM].

• Clarifies in IBC 2303.2.4 that FRTW must have the original product grade stamp in addition to the fire-retardant treatment labeling [S265-16]

ASCE 7-16 Minimum Design Loads and Associated Criteria for Buildings and Other Structures IBC changes regarding ASCE 7-16 are likely to lead to some confusion for designers and code officials. While the purpose of this article is to outline IBC changes that were approved, a few instances where changes were defeated are specifically noted below to allow for discussion of code paths to compliance with load provisions.

• Updates reference to ASCE 7-16 [ADM94-16] • Updates IBC wind and seismic load provisions to agree with updated criteria in ASCE 7-16 [S56-16 AM and S114-16 AM] • Updates references in Chapter 18 seismic provisions to coordinate with ASCE 7-16 [S166-16] continued on next page

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• A proposal to update IBC snow load provisions to agree with updated criteria in ASCE 7-16 was defeated during online voting [S103-16 AMPC1 Defeated]. Therefore, the IBC allows three paths rather than one: 1) Use the IBC to determine snow loads (Figure 1608.2 or Table 1608.2 for Alaska), 2) Use IBC 1608.2, which references ASCE 7-16, and use the new tables for the western US and New Hampshire, or 3) Use IBC 1608.2 to go to ASCE 7-16 which states that if an area is not in the new tables or exceeds the elevation limit (still a case study area), to then reference state produced maps which have greater detail for the western US and New Hampshire.

• A proposal removing LRFD and ASD load combinations based on reference to ASCE 7-16 was also defeated during online voting [S78-16 AM AMPC1 Defeated].

• On-line voting disapproved reference to ASCE 7-16 in IBC 1611 [S110-16 AMPC1 Defeated]. This has created a difference in requirements for secondary drains. ASCE 7-16 bases minimum requirements on a 15 min/100-yr event. IBC still uses the 1 hr/100-yr event for both primary and secondary minimum drain flow.

• Adds direct code references in IBC 803.11 to ASTM E 2579-13 and E 2404 for the mounting of laminate products, facings, and veneers with a wood substrate during testing [FS135-15 AS and FS136-15 AS] • Revises IBC 803.3 to require that cross-laminated timber and heavy timber elements be subject to the normal flame spread limitations for exits similar to other materials in exit enclosures; a previous exception for heavy timber elements within exit enclosures is inappropriate for exposed mass timber elements that make up entire wall and ceiling sections [FS132-15 AS] • Clarifies that fire officials can require roundthe-clock fire watch for construction that exceeds 40 feet above grade [F329-16] • Cleans up language permitting the use of FRTW sheathing in exterior walls of Type III and IV construction which is sometimes misinterpreted [G175-15 AS] • Releases FRTW and CLT exterior walls from having an assembly minimum thickness in favor of simply requiring a minimum actual thickness solely for the CLT [G18415 AS]; includes errata to the 2018 IBC as follows (errata shown “as written” using strikethrough/underline text):

Fire Protection

Fire-retardant-treated wood framing complying with Section 2303.2 shall be permitted within exterior wall assemblies not less than 6 inches in thickness with a 2-hour rating or less.

• Clarifies in IBC 704 that the protection of “gang studs” and built-up columns in the walls of lightweight construction can be provided by the membranes of the rated walls in which they are located [FS7-15 AM] • Reduces existing requirements in Table 705.2 on the location of building projections, such as roof overhangs [FS13-15 AS] • Relocates Chapter 14 Exterior Walls firerelated provisions for balconies, projections, and bay and oriel windows to Chapter 7 Fire and Smoke Protection Features [FS15-15 AM] • Revises IBC 706.2 to allow 3⁄4-inch plywood to run continuously through double firewalls in high seismic areas (Seismic Design Categories D and F) [FS29-15 AMPC1] • Clarifies sprinkler, fire partition, and draftstopping requirements in IBC 708 and 718 for multifamily structures; one change gives clear criteria for the sprinkler protection of attics without draftstopping [FS42-15 AMPC1] • Corrects certain prescriptive fire-resistance rated I-joist assembly description errors in Table 721.1(3) [FS129-15 AS and FS130-15 AS]

602.4.1 Fire-retardant-treated wood in exterior walls.

602.4.2 Cross-laminated timber in exterior walls.

Cross-laminated timber complying with Section 2303.1.4 shall be permitted within exterior wall assemblies not less than 6 4 inches in thickness with a 2-hour rating or less, provided the exterior surface of the cross-laminated timber is protected by one the following: 1) Fire-retardant-treated wood sheathing complying with Section 2303.2 and not less than 15⁄32-inch thick; 2) Gypsum board not less than ½-inch thick; or 3) A noncombustible material.

Heavy Timber and Mass Timber • Makes clear that SCL should be considered equivalent to heavy timber and clarifies the appropriate distinctions between nominal, net finished, and actual dimensions

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for heavy timber, glulam, SCL, and CLT [G178-15 AS] • Reorganizes heavy timber provisions aiding clear application of Type IV (heavy timber) construction requirements while also providing for separate application of code provisions that allow or specify the use of “heavy timber” elements outside of Type IV construction; moves certain heavy timber provisions from Chapter 6 to Chapter 23 and introduces a new table of minimum dimensions based on location within the building structure and loading condition [G179-15 AS; G180-15 AS] (Due to the extensive nature of these changes, the strikethrough/underline format is not shown in Appendix A. However, a summary of relocated sections is shown. The changes can be viewed on the ICC website.)

Construction Type • Permits the use of roofs for various occupancies without classifying the building as one containing an additional story, thus assuring continued flexibility for buildings of wood construction types [G24-15 AMPC2] • Permits performance-based alternatives for sound transmission design of floor assemblies using comparative engineering analysis [G190-15 AS]

Conclusion The 2018 IBC and 2018 IEBC are both available from ICC (www.iccsafe.org) and represent the state-of-the-art codes for design and construction of buildings outside the scope of the International Residential Code. These codes reference the latest wood standards such as the 2018 NDS and include other important changes to requirements for wood construction. In some situations, a building designer may want to use a more current code provision or consensus standard that is recognized in the building code adopted by a jurisdiction. In those cases, building officials, in accordance with Section 104.11 of the International Building Code, are permitted to accept designs prepared in accordance with newer consensus reference standards.

IBC 104.11 allows a jurisdiction to approve new technologies in materials and building construction, provided documentation supplied to the jurisdiction is found to assure equivalency in quality, strength, durability, and safety.▪

award winners and outstanding projects

Spotlight

Connecting Chicago’s “Second Shoreline”

By Kurt J. Naus, P.E., S.E., Matthew F. Hellenthal, P.E., S.E., and Daniel M. Gross, P.E. Alfred Benesch & Company was an Award Winner for its Chicago Riverwalk project in the 2017 Annual Excellence in Structural Engineering Awards Program in the Category – Other Structures.

F

or over 20 years, the City of Chicago had planned to enhance the main branch of the Chicago River by creating a continuous Riverwalk connecting the Lakefront to Chicago’s West Loop. Although the area has been referred to as the Riverwalk for many years, the path was interrupted by each crossing street and was only accessible via stairs at the ends of each block. The site was relatively featureless, with a concrete walking surface approximately seven feet above the water and only a few park benches, providing very little attraction to a very valuable piece of land along the river. During the reconstruction of the adjacent Wacker Drive Viaduct in the early 2000s, the roadway and structure were configured for future expansion. Sometime after, a renewed vision by the City aimed to develop the Riverwalk as Chicago’s “second shoreline” – a premier public space that would enable the city’s residents, workers, and visitors to connect with nature and the many recreational amenities offered by the urban waterway.

Designing and constructing the Riverwalk required extensive mitigation of existing structures below and above the water’s surface. To protect the many adjacent subsurface utilities, sewer outfalls, chiller pipes, freight tunnels, and the Chicago Transit Authority subway tunnels, extensive and sophisticated structure modeling was required to understand and monitor potential impacts to the existing infrastructure.

Above the water, creating the much-needed connections between the existing Riverwalk sections while preserving the integrity of the river’s historic bascule bridges was both an engineering and aesthetic challenge. The project team’s unique solution: canopied piers, or underbridges, were designed to create a new, continuous path along the six-block Riverwalk, which could resist substantial vessel collision loading and had minimal impact to existing historic structures. This critical design was achieved using drilled shafts that extended to the hardpan layer about 70 feet below the water’s surface. The structural framework was completed using an ingenious “bathtub” approach in which precast materials were delivered via barge and installed over the top of the drilled shafts. Concrete was

then placed into the precast form to complete each underbridge. This method minimized construction duration, cost, and disturbance in the river by eliminating the need for underwater formwork.

The interior shafts at each underbridge are located beneath the bascule bridges and required bridge lifts for installation. Test lifts were performed during the design phase of the project to determine the maximum opening angle that could be achieved. Existing plans were then used to develop clearance diagrams to aid the Contractor during the bid phase. Temporary closures of approximately one week were used at each street for installation.

Early Value Planning studies considered numerous alternatives for designing and constructing each of the six new “rooms” spanning the Riverwalk, ranging from pilesupported piers to floating docks. The desired and ultimately chosen design favored conventional tied back sheet pile walls to tie back to the existing wall where a concrete anchor block was cast over the wale, and timber piles stabilizing the new system, using existing tie rods. By incorporating the existing system in the final design, the number of new materials required and the carbon footprint were significantly reduced.

Previous construction of other Riverwalk segments revealed that the tie rods were in excellent condition with minimal corrosion when the concrete encasement was removed.

While there were not many concerns over the condition of the existing materials, consideration was still given to the redundancy of the system. In the event of an existing tie rod failure, the concrete anchor block and structural slab were designed and detailed with enough strength to redistribute the load to the adjacent tie rods. The structural slab is also able to transfer loads out to the peninsulas at each end of the rooms where additional lateral resistance can be developed.

This cost-effective solution made practical use of rehabilitating existing materials, reducing cost and construction time.

Because the Chicago River is navigable, a channel opening of 95 feet had to be

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Courtesy of christian phillips photography

maintained for boat traffic at all times during construction. Material deliveries were carefully maneuvered to the construction site by barges, and the majority of the work was completed off multiple barges adjacent to the dock wall at each section of the Riverwalk.

The contractor also had to consider the river’s ability to rise considerably following heavy rainfall. They implemented an overnight staging plan placing barges away from the historic bascule bridges so that, if waters were to rise, the bridges would not be damaged by construction equipment.

2,800 feet of new continuous walkway has transformed Chicago’s 1.3-mile Chicago Riverwalk, connecting the Lakefront to Chicago’s West Loop. The project’s six uniquely designed sections successfully integrated the original Riverwalk walls and utilized innovative precast underbridges supported on drilled shafts passing beneath each of the Chicago River’s historic bascule bridges.

Now complete and open to the public, the Riverwalk not only enhances the environment but enables residents, workers, and visitors to connect with the many recreational, cultural, and economic amenities offered by the Chicago River.▪ Kurt J. Naus is a Project Manager for Alfred Benesch & Company of Chicago. He can be reached at knaus@benesch.com.

Matthew F. Hellenthal is a Project Engineer for Alfred Benesch & Company of Chicago. He can be reached at mhellenthal@benesch.com.

Daniel M. Gross is a Senior Vice President and Construction Management Services Director for Alfred Benesch & Company of Chicago.

He can be reached at dgross@benesch.com.

NCSEA President’s Message: Demonstrate Our Passion

NCSEA News

News form the National Council of Structural Engineers Associations

Williston “Bill” Warren IV, P.E., S.E., SECB, F. SEAOC, NCSEA President

As we enter this new year, we accept new challenges and create new goals. Having previously held the President position of the Structural Engineers Association of California (SEAOC), I am familiar with the demands placed on non-profits and how to address these going into a new year. The focus of this next year will be in four areas: • Outreach and Advocacy • Increasing NCSEA’s partnerships • Bettering the Structural Engineering Summit • Demonstrating the passion Structural Engineers have for their profession as well as sharing the value Examining these areas one by one, a common goal stands out – growth – which lends plenty of opportunities for NCSEA and the SEA membership to continue succeeding together. To achieve this, we all must take part.

Following the footsteps of many SEAs that have practiced outreach and advocacy to successfully connect with the media, architects, building officials, and government officials, NCSEA will lend its national voice to the effort of increasing the perceived value the structural engineering profession brings to society. The general public routinely takes advantage of the fruits of the profession, buildings and others structures, without so much as a thought. Helping them to understand structural engineers’ contributions to society’s safety, health, and welfare will allow the profession deserved recognition.

Along with growing the knowledge of structural engineers’ contribution to general wellbeing, NCSEA also plans to continue to grow relationships with SEAs and various national organizations. With the constant restraint on our most important resource, time, NCSEA is looking to streamline efforts instead of duplicating or reinventing the wheel. Thanks to its 44 Member Organizations across the country, NCSEA is uniquely positioned to facilitate the distribution of ideas and products from one city, state, or region to another, ensuring the best is being utilized.

Increasing attendance at the 2018 Summit is already in my mind and those of the NCSEA board and staff. Having a record-breaking turnout at the 2017 event helped fuel recent conversations about bettering this year’s Summit. Along with increasing attendance, NCSEA wishes to continue fostering the celebration of the profession. Conferences like the NCSEA Structural Engineering Summit provide invaluable face-to-face networking opportunities that spark creativity and promote collaboration.

Growth as well as stability is at the core of every structural engineers calculations; these calculations ensure the safety, health, and welfare of the general public. In 2018, NCSEA wants structural engineers and society at large to take a step back to admire this profession’s accomplishments. It is time for structural engineers to proclaim their love for their profession and devote time to telling others what makes it important.

Taking the next step to complete these goals will require all of our participation; find your role and see it through.

Interviews just got easier! Find that perfect candidate or your next career move without leaving the comfort of your home or office. Visit the Career Center on www.ncsea.com for more information.

Emergency Response Training: CalOES SAP The California Office of Emergency Services (CalOES) Safety Assessment Program (SAP), hosted by NCSEA, is highly regarded as a standard throughout the country for engineer emergency responders. This course is one of only two post-disaster assessment programs that will be compliant with the requirements of the forthcoming Federal Resource Typing Standards for engineer emergency responders.

The training has been reviewed and approved by FEMA’s Office of Domestic Preparedness and provides engineers, architects and code-enforcement professionals with the basic skills required to perform safety assessments of structures following disasters.

This day long course consists of three webinar segments and is only offered live. Visit www.ncsea.com to register for the March 16, 2018 course.

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In an effort to improve and streamline the education process, NCSEA has introduced a new web home for its webinar program. By offering highquality webinars from expert speakers that are practical, accurate, and timely for the practicing structural engineer, NCSEA needed to improve the delivery system for those products with a new Education Portal.

The NCSEA Education Portal provides access to: • More than 20 high-quality Live Webinars per year.

• Recorded Webinar Library (100+ webinars), searchable by topic.

• SE Refresher and SE Exam Prep Course with more than 25 hours of first-rate instruction.

• Live CalOES training to help structural engineers become 2nd Responders in disaster situations. The next course is March 16, 2018.

• Your Continuing Education Certificates and Education History.

Visit www.ncsea.com for more information about the webinars coming up in 2018, the Webinar Subscription Plan, and to check out the new Education Portal!

NCSEA Webinars January 25, 2018 Multi-Family Wood Construction: Engineering Mid-Rise Buildings This presentation is intended for structural engineers who are seeking a full system understanding of the unique design considerations associated with 4-6 story wood structures. Structural design steps, considerations, and detailing best practices related to both gravity and lateral analysis will be covered.

Speaker: Ricky McLain, P.E., S.E. February 15, 2018 Structural Stability During Construction This course provides an overview of common areas of concern regarding structural stability during the construction process and will highlight issues that are often overlooked or missed during construction which can create unsafe conditions leading to partial or complete structural collapse.

Speaker: Matthew Pavelchak, P.E. February 27, 2018 Designing for Resilience: The Role of the Structural Engineer But what is “resilience,” and how will it affect structural engineering? In brief, resilience-based design shifts the emphasis from the safety of buildings to the recovery of communities from natural hazard events.

Speaker: David Bonowitz, S.E. Register at www.ncsea.com.

Courses award 1.5 hours of continuing education after the completion of a quiz.

Diamond Review approved in All 50 States.

Webinar time: 10:00 am Pacific, 11:00 am Mountain, 12:00 pm Central, 1:00 pm Eastern.

Booth Registration is Open Now–Reserve Your Space for 2018! Product Presentations are SOLD OUT STRUCTURE magazine

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News from the National Council of Structural Engineers Associations

In early January, NCSEA debuted the Education Portal, which handles all education purchases and makes it easier to find your education records to calculate PDHs. The Education Portal also has made it possible for NCSEA to improve the post-webinar certificate process. Certificates are now included with every Live Webinar registration for free and does not require additional registration. As always, multiple viewers in one location can attend webinars under one registration and receive credits after passing a quiz. This also has improved the function of NCSEA’s Webinar Subscription plan, which gives each subscriber unlimited access to all live NCSEA webinars, full access to the NCSEA Recorded Webinar Library 24/7/365, and an unlimited number of free CE certificates.

NCSEA News

NCSEA Rolls Out Education Portal to Enhance User Experience

Learning / Networking

Structural Columns

The Newsletter of the Structural Engineering Institute of ASCE

SEI/ASCE Live Webinars Learn from the Experts

January 9 – Practical Design of Bolted and Welded Steel Connections January 12 – Tornado Design Using ASCE 7-16 Commentary January 19 – Renovation of Pre-Engineered Buildings Individual Certificate Fee Discontinued.

Register at Mylearning.asce.org for these and much more.

Check out P.E./S.E. Exam Review Courses – www.asce.org/live_exam_reviews.

Structures Congress 2018 Register now and plan your participation using the new interactive planner at www.structurescongress.org.

Celebrate the Future of Structural Engineering at the Friday Reception hosted by CSI! Make sure to include your ticket when you register – available until April 13 or until sold out.

100% of ticket proceeds fund SEI strategic initiatives through the SEI Futures Fund in partnership with the ASCE Foundation.

Membership

SEI Elite Sustaining Organization Members

SEI Elite Sustaining Organization Members enjoy complimentary participation in the SEI Student Career Networking Event, April 20 at Structures Congress 2018 in Ft. Worth.

Reach more than 30,000 SEI members year-round with SEI Sustaining Organization Membership. Show your support for SEI to advance and serve the structural engineering profession. Learn more and join today at www.asce.org/SEI-Sustaining-Org-Membership.

Join or Renew SEI/ASCE Membership For innovative solutions and learning, to connect with leaders and colleagues, and enjoy member benefits: • Technical and professional news through SEI Update, STRUCTURE, Modern Steel Construction, and Civil Engineering Magazine • Member rates on SEI/ASCE conferences, continuing education, and publications • ASCE benefits, including small firm general and professional liability insurance, and 5 Free PDHs/yr • Get involved in an SEI Committee or Chapter effort Learn more and join or renew at www.asce.org/SEI.

Advancing the Profession

Structural Engineering Licensure Coalition Check out presentations, case studies, and resources at www.selicensure.org.

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CALLING ALL INNOVATORS, ENTREPRENEURS,

Expand your Horizons – Apply for Student and Young Professional Scholarships to Structures Congress

AND THE BEST MINDS OF OUR INDUSTRY:

Participating in Structures Congress opens doors to many SEI opportunities including meeting leaders of the profession, the latest knowledge from research, standards, and business practices, and fun networking events.

Encourage a student and/or young professional to join you at Structures Congress in Ft. Worth.

Last Call – Apply by January 5 at www.asce.org/SEI Made possible by the SEI Futures Fund in partnership with the ASCE Foundation.

The 2018 Innovation Contest is open for submissions!

Help to reshape the future of our nation’s infrastructure.

Learn more at www.asce.org/grand-challenge

SEI Student Career Networking Event April 20 at Structures Congress in Ft. Worth

Employers: Sign up now to participate, be included in event promotions, and receive student profiles in advance of the event. www.asce.org/SEI-Sustaining-Org-Membership Students: Plan now to connect one-on-one with employers for SE positions and internships. Learn more at www.asce.org/SEI-Students

SEI Online

Contract Documents Reduce conflicts and avoid litigation – the Engineers Joint Contract Documents Committee (EJCDC) develops and updates fair and balanced contract documents that represent the latest and best thinking in contractual relations among all parties involved in engineering design and construction projects. www.asce.org/contractdocuments.

SEI Welcomes Anne Ellis, P.E., F.ASCE

Check out more SEI News at www.asce.org/SEI including: SEI welcomes Anne Ellis, P.E., F.ASCE, to SEI Global Activities Executive Committee and SEI Futures Fund Board. Anne comments on her recent appointments, “Engineering achievement is never singular. It’s about teamwork and collaboration. The SEI Futures Fund provides our profession the means of investing collectively and collaboratively in those initiatives that will help us all to achieve.”

Errata SEI Standards Supplements and Errata including ASCE 7. See www.asce.org/SEI-Errata. If you would like to submit errata, contact Jon Esslinger at jesslinger@asce.org.

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The Newsletter of the Structural Engineering Institute of ASCE

Welcome to the new SEI Graduate Student Chapter at Penn State University

Structural Columns

Students and Young Professionals

CASE in Point

The Newsletter of the Council of American Structural Engineers

CASE Practice Guidelines Currently Available CASE 962-C – Guidelines for International Building Code-Mandated Special Inspections and Tests and Quality Assurance The Guideline is an update of the previous 3rd Edition to bring it current with the requirements of the 2012 International Building Code. The Guideline describes the roles and responsibilities of the parties involved in the special inspection and testing process, how to prepare a special inspection and testing program, the necessary qualifications of the special inspectors, how to conduct the program, and who should pay for the special inspections and test. The Appendix contains sample forms for specifying special inspections and tests, and sample letters to be filed with code-enforcement agencies after the program is completed.

CASE 962-D – A Guideline Addressing Coordination and Completeness of Structural Construction Documents The guidelines presented in this document will assist not only the structural engineer of record (SER) but also everyone involved with building design and construction in improving the process by which the owner is provided with a successfully completed project. Their intent is to help the practicing structural engineer understand the importance of preparing coordinated and complete construction documents and to provide guidance and direction toward achieving that goal.

Currently, the coordination and completeness of Documents vary substantially within the structural engineering profession and among the various professional disciplines comprising the

design team. The SER’s goal should be meeting both the owner’s and the contractor’s needs by producing a complete and coordinated set of Documents. Owners and contractors generally understand that some changes to the Documents will occur, because they realize that no set of Documents is perfect. The SER must focus on completeness, coordination, constructability, and the reduction of errors in order to minimize potential changes.

CASE 962-E – Self-Study Guide for the Performance of Site Visits During Construction This is a guide intended for the younger engineer but will be useful for engineers of all experience levels. Structural engineers know that site visits are crucial construction phase services that help clarify and interpret the design for the contractor. Site visits are also opportunities to identify construction errors, defects, and design oversights that might otherwise go undetected. Engineers should include adequate construction phase services, as a part of their scope of services, to ensure the design intent is properly implemented.

In 2016, the document was updated to include key points to summarize each section, updated references and definitions, and details on current tools for conducting site visits.

A companion document is available and was also updated in 2016: CASE Tool 10-1: Site Visit Cards.

You can purchase these and the other CASE Risk Management Tools at www.acec.org/case/news/publications.

CASE Winter Planning Meeting February 1 – 2, 2018; Austin, TX

The agenda for the meeting is: Thursday – February 1 6:15pm – 7:15pm Coalition Welcome Reception Sponsored by the Small Firm Council Friday – February 2 7:30am – 8:30am Shared Breakfast 8:00am – 12:00pm CASE General / Toolkit Committee Meeting CASE Contracts Committee Meeting CASE Guidelines Committee Meeting CASE Programs & Communications Committee Meeting STRUCTURE magazine

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12:00pm – 2:00pm Joint Coalition Roundtable Lunch Moderator: Andy Rauch, BKBM Engineers 2:15pm – 6:00pm CASE General / Toolkit Committee Meeting CASE Contracts Committee Meeting CASE Guidelines Committee Meeting CASE Programs & Communications Committee Meeting 6:00pm – 6:15pm Committee Wrap-up Session If you are interested in attending the meeting or have any suggested topics/ideas from a firm perspective for the committees to pursue, please contact Heather Talbert at htalbert@acec.org.

January 2018

The CASE Risk Management Convocation will be held in conjunction with the Structures Congress in Fort Worth, TX, April 19 – 21, 2018. For more information and updates go to www.structurescongress.org.

The following CASE Convocation sessions are scheduled to take place on Friday, April 20: 9:30 am – 10:30 am Managing Design Professionals’ Risk in the Design and Construction of Property Line Building Structures Moderator: Benjamin M Cornelius, Leslie E. Robertson Associates, R.L.L.P. Speaker: Kriton A. Pantelidis, Esq., Welby, Brady & Greenblatt, LLP

11:00 am – 12:30 pm The Good and the Bad of Delegated Design: How to Work With/As a Specialty Structural Engineer Moderator / Speaker: Kevin Chamberlain, DeStefano & Chamberlain Inc. 1:30 pm – 3:30 pm Construction Dispute Resolution through Forensic Engineering Moderator/Speaker: Benjamin M Cornelius, Leslie E. Robertson Associates, R.L.L.P. 3:30 pm – 5:00 pm Tackling Today’s Business Practice Challenges – A Structural Engineering Roundtable Moderator: Corey Matsuoka P.E., SSFM International, Inc.

CASE in Point

CASE Risk Management Convocation in Fort Worth, TX

2018 Small Firm Council Winter Seminar Looking to grow your business? Who isn’t!

Whether it is how to market, what key firm positions to fill, how to organize the firm into teams, or when it is time to delegate more, it is your leadership that ultimately grows your business. That is why you will want to join other small firm leaders from around the county for an in-depth examination of three strategic agendas your firm should implement to become more successful and profitable.

Seminar topics will focus on: • Sustainability – Promoting on-the-job learning and growth, and seeking new ways to improve your management process and protocols • Serviceability – Creating client value and a superior client experience through firm innovation and thought leadership • Survivability – Focusing on consistently generating and investing profits to expand your firm’s influence into current and prospective target markets This seminar is for firm leadership tasked with making decisions, such as owners, principals, HR professionals, CEOs, CFOs.

Registration:

ACEC Coalition Members – $399 ACEC Members – $499 | Non-members – $599

Location:

Hilton Garden Inn Austin Downtown & Convention Center 500 North IH 35, Austin, Texas 78701 Phone: 877-782-9444 use group code: ACEC Special Rate – $189/night until January 4, 2018, or until block sells out

About the Speaker Mark Goodale is a co-founder of Morrissey Goodale. His breadth of experience includes organizational development, strategic planning, mergers and acquisitions, marketing, and executive search. He is a trusted advisor and coach to dozens of industry executives.

Before helping to establish Morrissey Goodale, Mark was the Corporate Strategic Marketing Manager at PBS&J (now a part of Atkins) where he was charged with improving and implementing progressive corporate initiatives geared to position the firm for successful, large-scale client pursuits.

Before that, he worked at ZweigWhite for over a decade and headed the firm’s strategic business planning and marketing business units.

Mark has authored numerous articles for industry magazines such as Civil Engineering Magazine, CE News, and Consulting Specifying Engineer. He has been quoted many times in various industry publications and newspapers and is featured in the Morrissey Goodale/Axium video series, Building High-Performance Organizations. Mark was also a frequent contributor to ZweigWhite’s publications and events, and authored The Healthcare Market for AEP and Environmental Consulting Firms, the first of ZweigWhite’s market intelligence reports.

Always a top-rated speaker, Mark delivers presentations around the country on a wide variety of management topics at AIA, ACEC, NSPE, CSE, and ZweigWhite events.

Mark received his MBA from the Sawyer School of Business at Suffolk University where he now teaches Business.

To register for the seminar: http://bit.ly/2z2Mbrn Questions? Call 202-682-4377 or email at htalbert@acec.org.

Follow ACEC Coalitions on Twitter – @ACECCoalitions.

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CASE is a part of the American Council of Engineering Companies

Leading Innovation, Fostering Growth: Essentials to Achieving a Sustainable, Profitable Business February 1 – 3, 2018; Austin, TX

Business Practices

business issues

Looking for a Job? Part 1: Three Critical Steps to Finding Your Next Job

By Jennifer Anderson

S

earching for a job can be a long and challenging process. Regardless if you are a recent college graduate or an experienced engineer, read on to learn about the three critical steps that will result in a more productive and helpful job search.

When you are looking for a new job, there are universal truths that apply to the search process, no matter your age, years of experience, or location. The search process typically includes these stages: • The application process • Candidate evaluation process with an interview(s), reference checking, skill evaluation, etc.

• Acceptance or rejection of an offer Most people think about these stages as the only parts of a job search, but these are actually the last stages of the process. The job search starts as early as when you recognize that you want to make a job change. From that moment, you are on a path to define a significant change in your life, one that will likely impact your immediate family, too. Do not take a job search lightly; there is much at stake.

When you are ready to make a job change, here are three crucial steps that will provide more clarity when you get to the application, interview and offer stages of the job search.

Step 1: Define your Goals Be clear about what you want to accomplish with your professional career. I ask my clients, “When you get to the end of your life, what do you want to be known for?” – a simple, yet significant question. Imagine a life behind you; what do you want to be known for in your professional roles? What would you want your friends and closest professional associates to say about you? It is difficult to be a different person at work and at home, so aim to uncover what you want to be known for in both scenarios. It will help you to feel and be more synchronized in all roles of your life. This is also known as your personal brand.

For example, I have a client who wants to be known for helping his peers have access to useful information. This may seem like a simple thing to be known for but, if you stop to think about it, the results are enormous. He has noticed that when co-workers hoard information, they put the firm and their clients at risks. Whereas, when he can help team members communicate with each

other and share information, the feeling of camaraderie and unity is stronger. So, he looks for roles where he can help build trusting relationships between and amongst peers. When he decided to make a job change, he focused on firms with an established culture of collaboration that would provide greater opportunities to work in roles involving open communication and information sharing. This focus enabled my client to find work he loves with a firm which is a good fit. Without this focus going into the job search, he may have found himself again dissatisfied working for yet another company that does not champion collaboration and information sharing.

My client’s personal brand is consistently present in his home life, where he looks for ways to be connected to his family and community by also sharing helpful information.

He loves to read, travel, and learn about different cultures. He often seeks ways to connect various people and learn from a broad swath of society. I hope that, through his example, you see how a simple focus – being clear about what you want to be known for – is helpful to your job search. Do not skip this first step; be clear about the personal brand you are developing in all areas of your life.

Step 2: Manage your Network Once you understand what you want to be known for, manage your network effectively for proper introductions to your ideal firm.

If you are job searching by application to online job posts, you will find it very difficult to move the ball down the field. Filling out an online application is usually one of the steps in applying for a job, but it is not the best way to begin your introduction to a new firm. Instead, identify firms that you are interested in interviewing with and then evaluate your network to see who has connections at that firm. Leverage your connections to get a personal introduction to a manager at the firm; the job search will progress much more effectively than relying on an application sitting in an online applicant tracking system.

Step 3: Research Take the time to research companies, managers, and projects. As you understand the focus

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of a firm, it will be easier to know if you are interested in working for the firm and help you to be confident that the company, their culture, managers, teams, and projects are a good match for your personal brand. This is not to suggest that you will stay with the company for the rest of your working career. That is unrealistic in today’s work culture, but it is still essential that your next employer is a good match for how you want to build your career. Each of the companies you work for are going to make an impact on your overall career.

With research and deeper digging to truly understand a company, you will be better poised to ask insightful questions during the interview process and decide if the company will support where you want your career to go – and ensure that you are choosing a company that will enable you to build your personal brand. Without the research, you are not going to stand out from the other candidates.

Conclusion Remember that, during your job search, you still must jump through the red tape of applying online, interviewing, and deciding to accept/reject an offer. You will be more effective in making a solid decision if you clarify what you want to be known for, network your way into a company, and thoroughly research the potential new employer.

A future article will cover specifics of how to better interview and research prospective companies. I hope that you take to heart the three critical steps of the job search; both you and your new employer need you to be clear, focused, and on target with your job change.

I wish you good luck and encourage you to be yourself through the job search process!▪ Born into a family of engineers but focusing on the people side of engineering, Jennifer Anderson (www.CareerCoachJen.com) has nearly 20 years helping companies hire and retain the right talent. She may be reached at jen@careercoachjen.com.