The Great ShakeOut Earthquake Drill is an annual opportunity for people in homes, schools and organizations to practice what to do during earthquakes and improve their preparedness. In a post I wrote last October about the Great ShakeOut, I reminisced about the first earthquake I had to stop, drop and cover for – the Livermore earthquake in January, 1980. This year got me thinking about how our evacuation drills work.
At Simpson Strong-Tie, we use the annual Great ShakeOut drill to practice our building evacuation procedures. Evacuation drills are simple in concept – alarms go off and you exit the building. We have volunteer safety wardens in different departments who confirm that everyone actually leaves their offices. There are always a few people who want to stay inside and finish up a blog post. Once the building is empty and we have all met up in the designated meeting area, we do a roll call and wait for the all-clear to get back to work.
Several years ago the alarms went off. While waiting for the drill to end, we were concerned to see fire fighters arrive and rush into the building. Realizing this was not a drill, there were some tense moments of waiting. The fire chief and our president eventually walked out of the building and our president was yelling for one of our engineers. Turns out the engineer (who shall remain nameless) was cooking a chicken for lunch. Yes, a whole chicken. The chicken didn’t make it – I’m not sure what the guilty engineer had for lunch afterwards. At least we received extra evacuation practice that year. We aren’t allowed to cook whole chickens in the kitchen anymore.
Simpson Strong-Tie is helping increase awareness about earthquake safety and encouraging our customers to participate in the Great ShakeOut, which takes place next Thursday on October 20. It’s the largest earthquake drill in the world. More than 43 million people around the world have already registered on the site.
On October 20, from noon to 2:00 p.m. (PST), earthquake preparedness experts from the Washington Emergency Management Division and FEMA will join scientists with the Washington Department of Natural Resources and the Pacific Northwest Seismic Network for a Reddit Ask Me Anything – an online Q&A. Our very own Emory Montague will be answering questions. The public is invited to ask questions here. (Just remember that this thread opens the day before the event and not sooner.)
Emory, ready to answer some seismic-related questions.
We’re also providing resources on how to retrofit homes and buildings, and have information for engineers here and for homeowners here.
Earthquake risk is not just a California issue. According to the USGS, structures in 42 of 50 states are at risk for seismic damage. As many of you know, we have done a considerable amount of earthquake research, and are committed to helping our customers build safer, stronger homes and buildings. We continue to conduct extensive testing at our state-of-the-art Tye Gilb lab in Stockton, California. We have also worked with the City of San Francisco to offer education and retrofit solutions to address their mandatory soft-story building retrofit ordinance and have created a section on our website to give building owners and engineers information to help them meet the requirements of the ordinance.
Last year, Tim Kaucher, our Southwestern regional Engineering Manager, wrote about the City of Los Angeles’s Seismic Safety Plan in this post. Since that time, the City of Los Angeles has put that plan into action by adopting mandatory retrofit ordinances for both soft-story buildings and non-ductile concrete buildings. Fortunately, California has not had a damaging earthquake for some time now. As a structural engineer, I find it encouraging to see government policy makers resist complacency and enact laws to promote public safety.
Participating in the Great ShakeOut Earthquake Drill is a small thing we can all do to make ourselves more prepared for an earthquake. If your office hasn’t signed up for the Great ShakeOut Earthquake Drill, we encourage you to visit shakeout.org and do so now.
In last week’s blog post, we introduced the Simpson Strong-Tie® Strong-Wall® Wood Shearwall. Let’s now take a step back and understand how we evaluate a prefabricated shear panel to begin with.
First, we start with the International Building Code (IBC) or applicable state or regional building code. We would be directed to ASCE7 to determine wind and seismic design requirements as applicable. In particular, this would entail determination of the seismic design coefficients, including the response modification factor, R, overstrength factor, Ωo, and deflection amplification factor, Cd, for the applicable seismic-force-resisting system. Then back to the IBC for the applicable building material: Chapter 23 covers Wood. Here, we would be referred to AWC’s Special Design Provisions for Wind and Seismic (SDPWS) if we’re designing a lateral-force-resisting system to resist wind and seismic forces using traditional site-built methods.Continue Reading
Editor’s Note: This is a republished blog post with an introduction by Jeff Ellis.
This is definitely an attention-grabbing headline! At the National Earthquake Conference in Long Beach on May 4, 2016, Dr. Thomas Jordan of the Southern California Earthquake Center gave a talk which ended with a summary statement that the San Andreas Fault is “locked, loaded and ready to go.”
The LA Times and other publications have followed up with articles based on that statement. Temblor is a mobile-friendly web app recently developed to inform homeowners of the likelihood of seismic shaking and damage based on their location and home construction. The app’s creators also offer a blog that provides insights into earthquakes and have writtene a post titled “Is the San Andreas ‘locked, loaded, and ready to go’?” This blog post delves a bit deeper to ascertain whether the San Andreas may indeed be poised for the “next great quake” and is certainly a compelling read. Drop, cover and hold on!
Volkan and I presented and exhibited Temblor at the National Earthquake Conference in Long Beach last week. Prof. Thomas Jordan, USC University Professor, William M. Keck Foundation Chair in Geological Sciences, and Director of the Southern California Earthquake Center (SCEC), gave the keynote address. Tom has not only led SCEC through fifteen years of sustained growth and achievement, but he’s also launched countless initiatives critical to earthquake science, such as the Uniform California Earthquake Rupture Forecasts (UCERF), and the international Collaboratory for Scientific Earthquake Predictability (CSEP), a rigorous independent protocol for testing earthquake forecasts and prediction hypotheses.
In his speech, Tom argued that to understand the full range and likelihood of future earthquakes and their associated shaking, we must make thousands if not millions of 3D simulations. To do this we need to use the next generation of super-computers—because the current generation is too slow! The shaking can be dramatically amplified in sedimentary basins and when seismic waves bounce off deep layers, features absent or muted in current methods. This matters, because these probabilistic hazard assessments form the basis for building construction codes, mandatory retrofit ordinances, and quake insurance premiums. The recent Uniform California Earthquake Rupture Forecast Ver. 3 (Field et al., 2014) makes some strides in this direction. And coming on strong are earthquake simulators such as RSQsim (Dieterich and Richards-Dinger, 2010) that generate thousands of ruptures from a set of physical laws rather than assumed slip and rupture propagation. Equally important are CyberShake models (Graves et al., 2011) of individual scenario earthquakes with realistic basins and layers.
But what really caught the attention of the media—and the public—was just one slide
Tom closed by making the argument that the San Andreas is, in his words, “locked, loaded, and ready to go.” That got our attention. And he made this case by showing one slide. Here it is, photographed by the LA Times and included in a Times article by Rong-Gong Lin II that quickly went viral.
Believe it or not, Tom was not suggesting there is a gun pointed at our heads. ’Locked’ in seismic parlance means a fault is not freely slipping; ‘loaded’ means that sufficient stress has been reached to overcome the friction that keeps it locked. Tom argued that the San Andreas system accommodates 50 mm/yr (2 in/yr) of plate motion, and so with about 5 m (16 ft) of average slip in great quakes, the fault should produce about one such event a century. Despite that, the time since the last great quake (“open intervals” in the slide) along the 1,000 km-long (600 mi) fault are all longer, and one is three times longer. This is what he means by “ready to go.” Of course, a Mw=7.7 San Andreas event did strike a little over a century ago in 1906, but Tom seemed to be arguing that we should get one quake per century along every section, or at least on the San Andreas.
Could it be this simple?
Now, if things were so obvious, we wouldn’t need supercomputers to forecast quakes. In a sense, Tom’s wake-up call contradicted—or at least short-circuited—the case he so eloquently made in the body of his talk for building a vast inventory of plausible quakes in order to divine the future. But putting that aside, is he right about the San Andreas being ready to go?
Because many misaligned, discontinuous, and bent faults accommodate the broad North America-Pacific plate boundary, the slip rate of the San Andreas is generally about half of the plate rate. Where the San Andreas is isolated and parallel to the plate motion, its slip rate is about 2/3 the plate rate, or 34 mm/yr, but where there are nearby parallel faults, such as the Hayward fault in the Bay Area or the San Jacinto in SoCal, its rate drops to about 1/3 the plate rate, or 17 mm/yr. This means that the time needed to store enough stress to trigger the next quake should not—and perhaps cannot—be uniform. So, here’s how things look to me:
The San Andreas (blue) is only the most prominent element of the 350 km (200 mi) wide plate boundary. Because ruptures do not repeat—either in their slip or their inter-event time—it’s essential to emphasize that these assessments are crude. Further, the uncertainties shown here reflect only the variation in slip rate along the fault. The rates are from Parsons et al. (2014), the 1857 and 1906 average slip are from Sieh (1978) and Song et al. (2008) respectively. The 1812 slip is a model by Lozos (2016), and the 1690 slip is simply a default estimate.
So, how about ‘locked, generally loaded, with some sections perhaps ready to go’
When I repeat Tom’s assessment in the accompanying map and table, I get a more nuanced answer. Even though the time since the last great quake along the southernmost San Andreas is longest, the slip rate there is lowest, and so this section may or may not have accumulated sufficient stress to rupture. And if it were ready to go, why didn’t it rupture in 2010, when the surface waves of the Mw=7.2 El Major-Cucapah quake just across the Mexican border enveloped and jostled that section? The strongest case can be made for a large quakeoverlapping the site of the Great 1857 Mw=7.8 Ft. Teton quake, largely because of the uniformly high San Andreas slip rate there. But this section undergoes a 40° bend (near the ‘1857’ in the map), which means that the stresses cannot be everywhere optimally aligned for failure: it is “locked” not just by friction but by geometry.
A reality check from Turkey
Sometimes simplicity is a tantalizing mirage, so it’s useful to look at the San Andreas’ twin sister in Turkey: the North Anatolian fault. Both right-lateral faults have about the same slip rate, length, straightness, and range of quake sizes; they both even have a creeping section near their midpoint. But the masterful work of Nicolas Ambraseys, who devoured contemporary historical accounts along the spice and trade routes of Anatolia to glean the record of great quakes (Nick could read 14 languages!) affords us a much longer look than we have of the San Andreas.
The idea that the duration of the open interval can foretell what will happen next loses its luster on the North Anatolian fault because it’s inter-event times, as well as the quake sizes and locations, are so variable. If this 50% variability applied to the San Andreas, no sections could be fairly described as ‘overdue’ today. Tom did not use this term, but others have. We should, then, reserve ‘overdue’ for an open interval more than twice the expected inter-event time.
This figure of North Anatolian fault quakes is from Stein et al. (1997), updated for the 1999 Mw=7.6 Izmit quake, with the white arrows giving the direction of cascading quakes. Even though 1939-1999 saw nearly the entire 1,000 km long fault rupture in a largely western falling-domino sequence, the earlier record is quite different. When we examined the inter-event times (the time between quakes at each point along the fault), we found it to be 450±220 years. Not only was the variation great—50% of the time between quakes—but the propagation direction was also variable.
However, another San Andreas look-alike, the Alpine Fault in New Zealand, has a record of more regular earthquakes, with an inter-event variability of 33% for the past 24 prehistoric quakes (Berryman et al., 2012). But the Alpine fault is straighter and more isolated than the San Andreas and North Anatolian faults, and so earthquakes on adjacent faults do not add or subtract stress from it. And even though the 31 mm/yr slip rate on the southern Alpine Fault is similar to the San Andreas, the mean inter-event time on the Alpine is longer than any of the San Andreas’ open intervals: 330 years. So, while it’s fascinating that there is a ‘metronome fault’ out there, the Alpine is probably not a good guidepost for the San Andreas.
If Tom’s slide is too simple, and mine is too equivocal, what’s the right answer?
I believe the best available answer is furnished by the latest California rupture model, UCERF3. Rather than looking only at the four San Andreas events, the team created hundreds of thousands of physically plausible ruptures on all 2,000 or so known faults. They found that the mean time between Mw≥7.7 shocks in California is about 106 years (they report an annual frequency of 9.4 x 10^-3 in Table 13 of Field et al., 2014; Mw=7.7 is about the size of the 1906 quake; 1857 was probably a Mw=7.8, and 1812 was probably Mw=7.5). In fact, this 106-year interval might even be the origin of Tom’s ‘once per century’ expectation since he is a UCERF3 author.
But these large events need not strike on the San Andreas, let alone on specific San Andreas sections, and there are a dozen faults capable of firing off quakes of this size in the state. While the probability is higher on the San Andreas than off, in 1872 we had a Mw=7.5-7.7 on the Owen’s Valley fault (Beanland and Clark, 1994). In the 200 years of historic records, the state has experienced up to three Mw≥7.7 events, in southern (1857) and eastern (1872), and northern (1906) California. This rate is consistent with, or perhaps even a little higher than, the long-term model average.
So, what’s the message
While the southern San Andreas is a likely candidate for the next great quake, ‘overdue’ would be over-reach, and there are many other fault sections that could rupture. But since the mean time between Mw≥7.7 California shocks is about 106 years, and we are 110 years downstream from the last one, we should all be prepared—even if we cannot be forewarned.
Sarah Beanland and Malcolm M. Clark (1994), The Owens Valley fault zone, eastern California, and surface faulting associated with the 1872 earthquake, U.S. Geol. Surv. Bulletin 1982, 29 p.
Kelvin R. Berryman, Ursula A. Cochran, Kate J. Clark, Glenn P. Biasi, Robert M. Langridge, Pilar Villamor (2012), Major Earthquakes Occur Regularly on an Isolated Plate Boundary Fault, Science, 336, 1690-1693, DOI: 10.1126/science.1218959
James H. Dietrich and Keith Richards-Dinger (2010), Earthquake recurrence in simulated fault systems, Pure Appl. Geophysics, 167, 1087-1104, DOI: 10.1007/s00024-010-0094-0.
Edward H. (Ned) Field, R. J. Arrowsmith, G. P. Biasi, P. Bird, T. E. Dawson, K. R., Felzer, D. D. Jackson, J. M. Johnson, T. H. Jordan, C. Madden, et al.(2014). Uniform California earthquake rupture forecast, version 3 (UCERF3)—The time-independent model, Bull. Seismol. Soc. Am.104, 1122–1180, doi: 10.1785/0120130164.
Robert Graves, Thomas H. Jordan, Scott Callaghan, Ewa Deelman, Edward Field, Gideon Juve, Carl Kesselman, Philip Maechling, Gaurang Mehta, Kevin Milner, David Okaya, Patrick Small, Karan Vahi (2011), CyberShake: A Physics-Based Seismic Hazard Model for Southern California, Pure Appl. Geophysics, 168, 367-381, DOI: 10.1007/s00024-010-0161-6.
Julian C. Lozos (2016), A case for historical joint rupture of the San Andreas and San Jacinto faults, Science Advances, 2, doi: 10.1126/sciadv.1500621.
Tom Parsons, K. M. Johnson, P. Bird, J.M. Bormann, T.E. Dawson, E.H. Field, W.C. Hammond, T.A. Herring, R. McCarey, Z.-K. Shen, W.R. Thatcher, R.J. Weldon II, and Y. Zeng, Appendix C—Deformation models for UCERF3, USGS Open-File Rep. 2013–1165, 66 pp.
Seok Goo Song, Gregory C. Beroza and Paul Segall (2008), A Unified Source Model for the 1906 San Francisco Earthquake, Bull. Seismol. Soc. Amer., 98, 823-831, doi: 10.1785/0120060402
Kerry E. Sieh (1978), Slip along the San Andreas fault associated with the great 1857 earthquake, Bull. Seismol. Soc. Am.,68, 1421-1448.
Ross S. Stein, Aykut A. Barka, and James H. Dieterich (1997), Progressive failure on the North Anatolian fault since 1939 by earthquake stress triggering, Geophys. J. Int., 128, 594-604, 1997, 10.1111/j.1365-246X.1997.tb05321.x
Some contractors and framers have large hands, which can pose a challenge for them when they’re trying to install the holdown nuts used to attach our Strong-Wall® SB (SWSB) Shearwall product to the foundation. Couple that challenge with the fact that anchorage attachment can only be achieved from the edges of the SWSB panel, and variable site-built framing conditions can limit access depending upon the installation sequence. To alleviate anchorage accessibility issues, we’ve required a gap between the existing adjacent framing and SWSB panel equal to the width of a 2x stud to provide access so the holdown nut can be tightened. Even so, try telling a framer an inch and a half is plenty of room in which to install the nut!Continue Reading
If you’re one of the many engineers still confused by the ACI 318 – 11 Appendix D design provisions, this blog will help explain what’s required to achieve a ductile performing anchorage. Most building codes currently reference ACI 318 – 11 Appendix D as the required provision for designing a wide variety of anchor types that include expansion, undercut, adhesive and cast-in-place anchors in concrete base materials. This blog post will focus on section D.3.3.4.3(a) for an anchor located in a high seismic region. We’ll go over what these requirements are with a simple design example.
Ductility is a benefit in seismic design. A ductile anchor system is one that exhibits a meaningful degree of deformation before failure occurs. However, ductility is distinct from an equally important dimension called strength. Add strength, and a ductile steel element like the one shown in Figure 1 can now exhibit toughness. During a serious earthquake, a structural system with appreciable toughness (i.e., one that possesses both strength and ductility in sufficient degree) can be expected to absorb a tremendous amount of energy as the material plastically deforms and increases the likelihood that an outright failure won’t occur. Any visible deformations could help determine if repair is necessary.
Figure 1 – ½” mild steel threaded rod tensilely loaded to failure (starting stretch length = 8d)
Let’s start off with a simple example that will cover the essential requirements for achieving ductility and applies to any type of structural anchor used in concrete. We’ll arbitrarily choose a post-installed adhesive anchor. This type of anchor is very common in concrete construction and is used for making structural and nonstructural connections that include anchorage of sill plates and holdowns for shear walls, equipment, racks, architectural/mechanical/electrical components and, very frequently, rebar dowels for making section enlargements. We’ll assume the anchor is limited to resisting earthquake loading in tension only and is in seismic design category C – F. Section D.3.3.4.2 requires that if the strength-level earthquake force exceeds 20% of the total factored load, that the anchor be designed in accordance with section D.3.3.4.3 and D.3.3.4.4. We will focus on achieving the ductility option, (a), of D.3.3.4.3.
To understand anchor ductility we need to first identify the possible failure modes of an anchor. Figure 2 shows the three types of failure modes we can expect for an adhesive anchor located away from a free edge. These three failure modes generically apply to virtually any type of anchor (expansion, screw, cast-in-place or undercut). Breakout (Nb) and pullout (Na) are not considered ductile failure modes. Breakout failure (Nb) can occur very suddenly and behaves mostly linear elastic and consequently absorbs a relatively small amount of energy. After pullout failure (Na) has been initiated, the load/displacement behavior of the anchor can be unpredictable, and furthermore, no reliable mechanism exists for plastic deformation to take place. So we’re left with steel (Nsa). To achieve ductility, not only does the steel need to be made of a ductile material but the steel must govern out of the three failure modes. Additionally, the anchor system must be designed so that steel failure governs by a comfortable margin. Breakout and pullout can never control while the steel yields and plastically deforms. This is what is meant by meeting the ductility requirements of Appendix D.
Figure 2 – Three possible failure modes for an adhesive anchor loaded in tension
Getting back to our design example, we have a single post-installed 5/8” diameter ASTM F1554 Gr. 36 threaded rod that’s embedded 12” deep, in a dry hole, in a concrete element that has a compressive strength of 2,500 psi. The concrete is 18” thick and we assume that the edge distance is large enough to be irrelevant. For this size anchor, the published characteristic bond strength is 743 psi. Anchor software calculations will produce the following information:
The governing design strength is compared to a demand or load combination that’s defined elsewhere in the code.
Here’s the question: Before proceeding with the remainder of this blog, judging by the design strength values shown above, should we consider this anchorage ductile? Your intuition might tell you that it’s not ductile. Why? Pullout clearly governs (i.e., steel does not). So it might come as a surprise to learn that this adhesive anchor actually is ductile!
To understand why, we need to look at the nominal strength (not the design strength) of the different anchor failure modes. But first let’s examine the equations used to determine the design strength values above:
The above values incorporate the notation φ (“phi”) and a mandatory 0.75 reduction factor for nonductile failure modes (Ncb ,Na) for applications located in high seismic areas (seismic design category C–F). The φ factor is defined in section D.4. However, manufacturers will list factors specific to their adhesive based on anchor testing. The mandatory 0.75 reduction comes from section D.3.3.4.4 and is meant to account for any reduction associated with concrete damage during earthquake loading. The important thing to remember is that the nominal strength provides a better representation of the relative capacity of the different failure modes. Remove these reduction factors and we get the following:
Now steel governs since it has the lowest strength. But we’re not done yet. Section D.3.3.4.3.(a).1 of Appendix D requires that the expected steel strength be used in design when checking for ductility. This is done by increasing the specified steel strength by 20%. This is to account for the fact that F1554 Gr. 36 threaded rod, for example, will probably have an ultimate tensile strength greater than the specified 58,000 psi. (Interestingly, the ultimate strength of the ½” threaded rod tested in Figure 1 is roughly 74 ksi, which is about 27% greater than 58,000 psi.) With this in mind, the next step would be to additionally meet section D.3.3.4.3.(a).2 such that the following is met:
By increasing the steel strength by 20%, the nominal strength of the nonductile failure modes (Ncb ,Na) must be at least that much greater to help ensure that a ductile anchor system can be achieved. The values to compare finally become:
Now steel governs, but one more thing is required. As shown in Figure 3, Section D.3.3.4.3.(a).3 of Appendix D also requires that the rod be made of ductile steel and have a stretch length of at least eight times the insert diameter (8d). Appendix D defines a ductile steel element as exhibiting an elongation of at least 14% and a reduction in area of at least 30%. ASTM F1554 meets this requirement for all three grades of steel (Grade 36, 55 and 105) with the exception of Grade 55 for anchor nominal sizes greater than 2”. Research has shown that a sufficient stretch length helps ensure that an anchor can experience significant yielding and plastic deformation during tensile loading. The threaded rod shown in Figure 1 was tested using a stretch length of 4” (8d). Lastly, section D.3.3.4.3.(a).4 requires that the anchor be engineered to protect against buckling.
Figure 3 – Stretch length
Appendix D doesn’t require that an anchor system behave ductilely. Three additional options exist for Designers in section D3.3.4.3. Option (b) allows for the design of an alternate failure mechanism that behaves ductilely. Designing a base plate (or support) that plastically hinges to exhibit ductile performance is one example. Option (c) involves a case where there’s a limit to how much load can be delivered to the anchor. Although option (c) under D.3.3.4.3 falls under the tensile loading section of Appendix D, the best example would apply to anchorage used to secure a wood sill plate or cold-formed steel track. We know from experiments that the wood crushes or the steel yields and locally buckles at a force less than the capacity of the concrete anchorage. Clearly energy is absorbed in the process. The most commonly used option is (d), which amplifies the earthquake load by Ωo. Ωo can be found in ASCE 7 – 10 for both structural and nonstructural components. The value of Ωo is typically taken to be equal to 2.5 (2.0 for storage racks) and is intended to make the anchor system behave linear elastically for the expected design-level earthquake demand.
These same options exist for shear loading cases. However, achieving system ductility through anchor steel is no longer an option for shear loading according to ACI 318 – 11, because the material probably won’t deform appreciably enough to be considered ductile.
While factors such as edge-distance and embedment-depth restrictions make achieving ductility difficult for post-installed anchors, it should come as some consolation that in many cases the Designer can achieve ductile performance for cast-in-place anchors loaded in tension through creative detailing of reinforcing steel (section D.5.2.9) to eliminate breakout as a possible failure mode. This has been explored in some detail in two previous Simpson Strong-Tie blogs titled “Anchor Reinforcement for Concrete Podium Slabs” and “Steel Strong Wall Footings Just Got a Little Slimmer.”
Have you ever been at home during an earthquake and the lights turned off due to a loss of power? Imagine what it would be like to be in a hospital on an operating table during an earthquake or for a ceiling to fall on you while you are lying on your hospital bed.
While the contents of this blog are certainly not what Abraham Lincoln had in mind when he made the statement that I’m using to title this blog post, it does speak volumes to the pertinence of what will be discussed today. “Design by others” or some variation of this appears in many parts of Simpson Strong-Tie details.Continue Reading
As soon as news spread that 7.8-magnitude and 7.3-magnitude earthquakes struck Nepal in April and May of this year, earthquake structural engineering experts from our firm, Miyamoto International, hopped on planes from three countries to offer assistance. We do this in hopes that our expertise and technical advice might help stricken communities recover; help them to build better and ultimately help save lives.Continue Reading
While 54 inches is a good height and will get you on most amusement park rides, what about this dimension for the width of a footing? We did some tests recently — actually a lot of tests — that answered that question.
Steel Strong-Wall® narrow panels are great for resisting high seismic or wind loads, but due to their narrow widths, their resulting anchor uplift forces can be rather hefty, requiring very large pad footings. How large? For Seismic Design Categories C-F, the largest cracked concrete solution per ACI318-11 Appendix D has a width of 54 inches and an effective embedment depth of 18 inches in order to ensure the anchor remains ductile. The overall length of this footing, as seen in Figure 1, can be up to 132 inches. While purely code driven, these solutions have historically presented challenges in the field. Most concrete contractors have to dig footings this size by hand. This often leads to discussions with their engineers about finding a better solution.
Simpson Strong-Tie has been studying cast-in-place anchorage extensively in recent years. Our research has been featured in a couple of blog posts: The Anchorage to Concrete Challenge – How Do You Meet It? and Podium Anchorage – Structure Magazine. Concrete podium slab anchorage was a multi-year test program that started with grant funding from the Structural Engineers Associations of Northern California for initial concept testing at Scientific Construction Laboratories Inc. and wrapped up with full-scale detailed testing completed at the Simpson Strong-Tie Tye Gilb Laboratory in Stockton, California. This joint venture studied the performance of anchorage reinforcement into thin podium deck slabs (10-14 inch) to resist the high overturning forces of continuous rod systems on 4-5-story mid-rise construction. The testing confirmed the need to comply with Appendix D requirements to prevent plastic hinging at anchor locations. Be on the lookout for an SE Blog post on that topic in the near future. Armed with what we learned, we decided to develop tested anchor reinforcing solutions for the Steel Strong-Wall.
The newly developed anchor reinforcement solutions for grade beams are calculated in accordance with ACI318 Appendix D and tested to validate performance. Anchor reinforcement isn’t a new concept, as it’s been in ACI318 for some time. Essentially, anchor reinforcements transfer load from the anchor bolt to the reinforcing, which restrains the breakout cone from occurring. For the new grade beam details, the additional ties near the anchor are designed to resist the load from the anchor and are developed into the grade beam. The new details offer solutions with widths as narrow as 18 inches when anchor reinforcement is used.
Two details have been developed: one for the larger panels (SSW18, SSW21, SSW24) as shown in Detail 1/SSW1.1, and one for the smaller panels (SSW12, SSW15) as shown in Detail 2/SSW1.1. The difference between the two is the number of anchor reinforcement ties specified in Detail 3/SSW1.1. For SSW18, SSW21 and SSW24 panels (Detail 1/SSW1.1), the total number of reinforcement per anchor is specified. Due to their smaller sizes, the anchor reinforcement ties specified in Detail 2/SSW1.1 for the SSW12 and SSW15 panels are the total required per panel.
Detail 1/SSW1.1Detail 2/SSW1.1Detail 3/SSW1.1
Validation Testing
From the concrete podium deck anchorage test program, we discovered that the flexural and shear capacity of the slab is critical to anchor performance and must be designed to exceed the demands created by the attached structure. For grade beams, this also holds true. In wind-load applications, this demand includes the factored demand from the Steel Strong-Wall. In seismic applications, our testing and analysis showed that achieving the anchor performance expected by Appendix D design methodologies requires the concrete member design strength to resist the amplified anchor design demand from Appendix D Section D.3.3.4.3.
Validation testing was conducted to evaluate this concept. The test program consisted of a number of specimens with different configurations, including:
Closed tie anchor reinforcement
Non-closed tie u-stirrup anchor reinforcement
Control specimen without anchor reinforcement
Flexural and shear reinforcement were designed to resist Appendix D amplified anchorage forces and were compared to test beams designed for non-amplified strength level forces. The results of the testing are shown in Figure 2. In the higher Seismic Design Categories (C-F), the anchor assembly must be designed to satisfy Section D.3.3.4.3 in ACI318-11 Appendix D. In accordance with D.3.3.4.3 (a), the concrete breakout strength needs to be greater than 1.2 times the nominal steel strength of the anchor, 1.2NSA. This requires a concrete breakout strength of 87 kips for a Steel Strong-Wall that uses a 1-inch high-strength anchor.
Figure 2: Steel Strong-Wall Grade Beam Testing
Grade beams without the anchor reinforcement detail and with flexural and shear reinforcement designed to the Appendix D amplified anchorage forces performed similar to those with closed-tie anchor reinforcement and flexural and shear reinforcements designed to the non-amplified strength level forces. Both, however, came up short of the necessary forces required by Section D.3.3.4.3 (a). From Test V852, we discovered that even though the flexural and shear reinforcement were designed with the amplified forces, the non-closed tie u-stirrups did not ensure the intended performance. From observation, the u-stirrups do not provide adequate confinement of the concrete and tend to open up under loading conditions, resulting in splitting of the beam at the top as can be seen in the photo.
Test V852: Non-Closed U-Stirrups
Tests W785 and W841 resulted in the best performance. Both test specimens contained flexural and shear reinforcement designed for the amplified forces, as well as closed-ties. Two configurations were tested to study their performance — two piece closed-tie anchor reinforcement in W785 and a single piece closed-tie anchor reinforcement in Test W841. As seen in Figure 2, their performance was very similar, and met the requirements of Section D.3.3.4.3 (a). The closed-ties helped confine the concrete near the top of the beam, allowing the assembly to reach the expected performance load (See the photo below). It’s important to indicate the following specifics in the New Grade Beam Anchor Reinforcement Details:
Anchor Reinforcement is #4 closed-ties
SSWAB embedment depth is 16″ +/- ½” (as shown in Detail 3/SSW1.1). This is to ensure there is enough development length of the anchor reinforcement on both sides of the theoretical breakout surface as required by ACI318-11 D.5.2.9.
The minimum distances from the anchor bolt plate washer to top and bottom of closed tie reinforcement are 13 inches and 5 inches respectively to ensure proper development above and below the concrete breakout cone (refer to Detail 3/SSW1.1).
The spacing between the two vertical legs of the anchor reinforcement tie must be 10 inches apart. While this may differ from your shear reinforcement elsewhere in the grade beam, it ensures the reinforcement is located close enough to the anchor and adequate development length is provided.
Flexural reinforcement (top and bottom) and shear reinforcement (ties throughout the grade beam length) are per the designer. Simpson Strong-Tie has provided information in Detail 3/SSW1.1 for the applicable minimum LRFD Applied Design Seismic Moment (See Figure 3) to make sure the grade beam design will at least resist the applied anchor forces. Project design loads not related to the Strong-Wall panel also should be considered and could control the grade beam design.
Closed-Tie Anchor ReinforcementFigure 3: LRFD Applied Design Seismic Moment
Simpson Strong-Tie is interested in hearing your thoughts on the new details. What is your opinion? How have the new details been received on your job sites?
This post is the second of a two-part series on the results of research on anchorage in reinforced brick. The research was done to shed light on what tensile values can be expected for adhesive anchors. In last week’s post, we covered the test set-up. This week, we’re taking a look at our results and findings.Continue Reading
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next generation of super-computers—because the current generation
overlapping the site of the Great 1857 Mw=7.8 Ft. Teton quake, largely because of the uniformly high San Andreas slip rate there. But this section undergoes a 40° bend (near the ‘1857’ in the map), which means that the stresses cannot be everywhere optimally aligned for failure: it is “locked” not just by friction but by geometry.
