Making Earthquakes Tremble

Carquinez Bridge designed to stand up to the next big seismic event

Bridges Article May 17, 2002
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Long-span suspension bridges with spans over 2,300 ft have
not been built in the U.S. for the past 36 years when the 4,260-ft main span
Verrazano Narrows bridge opened to traffic in 1964. In 2004, the Carquinez
Bridge, with a main span of 2,388 ft, will be ready for action. It will be the
first orthotropic steel box girder suspension bridge ever built in the country.

The bridge, which spans the Carquinez strait about 20 miles
northeast of San Francisco, is located within a few miles of several active
faults. While other famous suspension bridges like the Golden Gate Bridge and
the San Francisco-Oakland Bay Bridge are undergoing major seismic retrofit,
Carquinez Bridge is the first bridge in the U.S., located in a potentially high
seismic risk area, to be designed to present-day stringent seismic design
standards.

The first Carquinez bridge, designed by David Steinman and
Charles Derleth, was built in 1927 by the American Toll Bridge Co. It replaced
a heavily used ferry service and cleared a bottleneck on the highway between
San Francisco and Sacramento. The state of California acquired the bridge in
1941 and in the face of growing traffic needs designed and constructed a second
Carquinez Bridge in 1958. Completion of the second bridge was part of an
ambitious project that transformed a two-lane road into the modern eight-lane
interstate highway of today. Both bridges are multi-span cantilever truss
bridges with maximum spans of about 1,180 ft.

Considering the potential for seismic activity in the
Carquinez strait, safety initiatives prompted the California Department of
Transportation (Caltrans) to commission retrofit studies of the existing
structures. These studies were carried out by a joint venture team of HNTB,
Oakland, and CH2M Hill, Oakland. As a result of these studies, replacement over
retrofit was considered the best “overall value” for the 1927
bridge due to several reasons: the bridge had severe seismic deficiencies; its
main members required rehabilitation due to fatigue and corrosion; the concrete
deck required replacement; and the lane width were substandard with no
shoulders. The state chose to replace the 1927 bridge, and the 1958 bridge was
considered for seismic retrofit for continued service.

Caltrans chose the joint venture team of De Leuw, Cather
& Co., San Francisco, Steinman Boynton Gronquist & Birdsall, New York,
(both firms are now part of Parsons Transportation Group) and OPAC Consulting
Engineers, San Francisco, to design the replacement bridge. A type selection
study was undertaken by the consultants to choose between cable stayed and
suspension bridge alternatives. The study concluded that a two-tower suspension
bridge would cost about the same as a three-tower cable-stayed bridge. However,
for reasons of better seismic performance, shorter construction schedule,
better aesthetics and less construction risk, the suspension bridge alternative
was chosen for final design.

 

Everything ties into the towers

The new Carquinez Bridge has a total length of 3,464 ft
between the north anchorage and the transition pier. The bridge will carry four
lanes of traffic along I-80 in addition to a pedestrian/bike lane. The overall
width of the bridge deck between the main cables is 89.2 ft.

 

Towers: Two
reinforced concrete towers of the bridge will rise about 410 ft above the water
level. Each tower leg consists of a hollow box section with spirally reinforced
corner pilasters about 3.3 ft in diam. connected by 19.7-in.-thick walls. This
hollow cellular section was shown to provide superior seismic performance as
well as agreeable aesthetics. The tower legs have been designed to behave
elastically under the most severe design earthquake forces with limited
inelasticity being permitted at the tower base. Two reinforced concrete struts
connect the two tower legs. The struts have been designed to remain elastic.
The strut nominal capacity is 20% greater than the plastic moment capacity of
the tower legs. This design ensures that any inelasticity will occur in the
more ductile tower leg rather than the post-tensioned struts. Tower legs at
both the north and south towers are currently under construction using the jump
form method of construction.

 

Tower piles: Each
tower leg is supported on six 9.8-ft-diam. drilled shaft piles. Steel casings
with an average length of about 155 ft for the south tower and about 127 ft for
the north tower are driven to bedrock. The rock sockets are about 141 ft long
at the south tower and about 79 ft long at the north tower. The piles are
detailed to provide adequate ductility and confinement to withstand the large
seismic forces.

Tower foundations posed one of the biggest construction
challenges of the project. Construction of the north tower piles involved
driving the casings to the specified tip. The casings were then cleaned out and
an 8.9-ft-diam. rock socket was drilled beyond the tip of the casing. Polymer
slurry with 6% salt concentration was used to stabilize the hole and to
minimize softening and swelling of soft claystone and siltstone present in the
rock. Once the rock socket tip elevation was reached, the rebar cage was
lowered and concrete placed using a tremie pipe displacing the slurry. The
rebar cage was then lowered in the casing and lapped with the rock socket cage.
Concrete was then poured into the casing to complete the pile construction.

The rock mass at the south tower piles is highly fractured,
and polymer slurry was not able to stabilize the rock socket hole. Thus, a
different construction method was employed. Once the casing was in place, a
10.8-ft-diam. under-reamer was used to drill about 25-ft-deep segments, which
were stabilized using polymer slurry. A lean concrete mix was then placed by
tremie. Once the concrete hardened, an 8.9-ft-diam. hole was cored through the
concrete ensuring a stable hole for the rock socket. Thus, proceeding in
segments of about 25 ft, the rock socket tip was reached. Pile construction was
then completed using the same procedure that was used at the north tower piles.

 

Tower foundations and footing forms: style='font-weight:normal'> The tower foundations are about 7.6 ft below the
mean sea level. To avoid construction of the footings under water, the
designers envisioned the use of precast footing forms. These precast forms were
supposed to be floated to the site and positioned in place at the exact
location and elevation over an erection frame supported on temporary piles. The
precast form would act as a cofferdam for the construction of the footing and
also act as a template for driving the drilled shafts.

However, the contractor chose to modify this construction
method. Since the construction of the drilled shafts was on the critical path,
the contractor started the construction of the shafts sooner while the precast
forms were still being fabricated. Once the drilled shafts were constructed and
the footing forms fabricated, the forms were floated in place at high tide and
lowered on top of the piles. The contractor used a positioning system for the
drilled shafts that allowed tight tolerances on the horizontal position of the
steel casings, which ensured proper mating of the precast footing forms to the
casings.

This construction technique will now be used on two other
San Francisco Bay-area bridges which use large diameter drilled shaft piles and
footings which are below water: the new Benicia Martinez Bridge and the San
Francisco-Oakland Bay Bridge east span replacement project.

 

Transition pier: The
pier is located at the south end of the suspension bridge where it transitions to
a steel girder supported span of the approach viaduct. The approach viaduct is
designed by Caltrans and the real challenge in designing the structure was to
mate the two bridges with vastly different dynamic response during an
earthquake. The suspension bridge is much more flexible than the approach
viaduct. Thus the steel box girder superstructure of the viaduct is designed as
an isolation structure to accommodate up to 4.9 ft of differential movement
longitudinally between the two bridges, providing a flexible link.

The pier column cross section is a hollow rectangular box
with spirally tied corner pilasters. The bent cap supports the suspension
bridge loads through tall rocker links and also loads from the approach viaduct
girders. In keeping with the Caltrans seismic design philosophy, the pier is
designed with a strong cap/weak column with plastic hinges forming at the
column top and bottom in a transverse earthquake.

In addition to supporting loads from the bridge decks, the
pier also supports cable tie-down loads. The tie-down sections pass through the
hollow columns and are anchored into the footings. To prevent the tie-down
forces from being transferred to the columns, tie-down anchor rods are grouted
in the solid sections at the base of the columns and the cap only after the
superstructure box girder is erected and its load transferred to the footing.
Each column is supported on 25- to 30-in. cast-in-steel shell concrete (CISS)
pile. Because of the high tensile load from the tie-down, the net compression
force on the footing is considerably reduced.

 

Deck: The new
Carquinez Bridge will be the first major suspension bridge in the U.S. to use a
closed-steel orthotropic box girder deck. This design is particularly effective
in resisting bending stresses because of the large flange area. At the same
time the closed box shape provides adequate torsional rigidity against wind,
seismic and other asymmetrical loads. In cross section, the box girder is 9.8
ft deep and 95.1 ft wide. The edge plate and side plates are shaped to provide
an aerodynamically stable cross section.

The suspended superstructure is designed as a continuous
element with expansion joints only at the extreme ends of the girder resulting
in a total girder length of 3,464 ft. The box girder offers many advantages
over truss-stiffened suspension spans including lower steel costs and lower
maintenance costs, since much of the steel surface is sheltered within the
section. The design of the deck incorporated several advances in fatigue endurance
and innovative use of groove-welded connections. These innovative modifications
were made and validated as part of full-size physical testing for the
Williamsburg bridge deck replacement in New York City.

Longitudinal bulkheads are provided to improve the vertical
shear carrying capacity of the box and to accommodate seismic demands resulting
in large bending moments about the vertical axis of the girder. The bulkheads
also act as longitudinal distributing members along the box girder between two
transverse bulkheads. Transverse bulkhead spacing is determined by the spacing
of the suspenders, which is 40.7 ft. Coincident with each suspender, a
transverse bulkhead is provided with an intermediate bulkhead between each of
them. A determining factor for spacing the transverse bulkheads is the maximum
desirable span of the orthotropic deck. By maximizing the span of the
orthotropic deck, the number of transverse bulkheads is reduced by more than
25%, as compared to the spacing of about 15 ft incorporated into several
European and Japanese designs.

Initial fabrication of the bridge deck is being carried out
by the Japanese firm Ishikawajima-Harima Heavy Industries Co. Ltd. at their
plant in Japan. The fabricated box girder segments will be transported across
the Pacific Ocean to the Carquinez bay and lifted directly onto the span in
their final position.

 

Suspension system:
Each main cable consists of 8,584 zinc-coated, carbon-steel wires compacted
into a diameter of 20.2 in. The No. 6 gage cold drawn wire has a specified
minimum ultimate strength of 228 ksi and a minimum yield strength by the 0.2%
offset method of 170 ksi. A maximum allowable working stress of 100 ksi was
used.

Following compaction of the main cable into its circular
shape and prior to wrapping it with No. 9 gage galvanized steel wire, the cable
will be coated with a high metallic zinc content polyurethane base
waterproofing paste.  This paste
product was successfully used on the Storebealt Bridge in Denmark and has
gained favor for several recently constructed suspension bridges in Norway.
Placed to fill the interstices between the wrapping wire and the main cable
wires, the paste will physically bond to the wire surfaces, providing a
flexible and waterproof barrier. Over-coating the wrapping wire with a
multi-coat system of highly elastic acrylic-polymer paint provides further
waterproofing.

In the splay chambers of the gravity anchorages the main
cables each splay into 37 individual strands, with each strand secured around
strand shoes of conventional design. This design approach considers that the
cables will be constructed by air spinning, which reflects the
contractor’s selected method of cable erection. Each strand shoe is
connected to a pair of anchor rods that extend to the rear of the gravity
anchor block and attach to a series of steel anchor girders embedded in the
concrete anchorage.

The suspenders are standard zinc-coated steel structural
wire rope. For the manufacture of the suspender assemblies, resin socketing has
been selected over the alternate zinc socketing for improved fatigue
performance and corrosion prevention. The suspenders are connected to the box
girder with external linkage assemblies providing a maintainable and
inspectable system. All linkage components are hot-dip galvanized with
stainless steel pins, making the entire assembly consistent with the specified
150-year design life. The short suspender assemblies near the expansion joints
at the end of the side spans have been provided with low-friction spherical
bearings to provide for multi-directional rotation of the socket attachments.
Additionally, each link has a lug for the attachment of a jacking device to
facilitate future suspender replacement.

Suspenders are designed with a safety factor of 4.0 against
the ultimate breaking strength, including efficiency losses resulting from
looping the ropes over the cable bands. 
A built-in redundancy is provided in the design by including a pair of
suspender ropes at each connection point. The design also provides sufficient
suspender capacity in the unlikely event of the loss of one suspender group.

 

Anchorages: The
subsurface condition at the two cable anchorages is significantly different. As
a result their geometry as well as their foundations are different.

 

North anchorage: This
is a gravity-type anchorage, which is located high up on a rock buff
overlooking the Carquinez strait. The anchorage is combined with the bridge
abutment, which supports the bridge deck through tall rocker bearings. The
underlying rocks are folded and fractured with the strike and dip varying
across the site. Geologic mapping and subsurface exploration identified a
series of shear zones and faults, though there was no evidence to suggest that
these features were of tectonic origin. To reduce the pressure on the rock
buff, the anchorage is located about 115 ft from the face.

The top of the buff will be benched to allow the bridge
girder to span to the abutment. The heel of anchor blocks for each cable is
embedded about 65 ft in the rock and rises in steps up to the abutment seat.
Resistance to the large cable forces is provided by base friction as well as
the passive resistance against the embedded structure. To enhance the
resistance of the anchor blocks against block failure of the buff the two
blocks are connected by a massive grade beam. This helps to mobilize a much
larger block of rock mass providing increased factor of safety against failure.

 

South anchorage:
Unlike the north anchorage, the south anchorage is located on fill underlain by
soft clay, loose sand and weathered rock. With high ground water, the sand was
found to be susceptible to liquefaction. The lateral resistance to the cable
pull is thus provided by 380-30-in.-diam. CISS piles—171 of these piles
are vertical and the rest are driven at a batter of 3:1, with the batter being
in the direction of the cable pull. The cable splay chambers and part of the
anchor blocks are above ground with the remaining anchor block embedded below
grade. The anchorage is located well south of the transition pier and below the
approach viaduct. Its location was partly determined by the presence of Union
Pacific railroad tracks, which were to be cleared by the main cables.

 

Strong on all levels

The bridge is being designed for two levels of earthquake,
the Safety Evaluation Earthquake (SEE) corresponding to a maximum credible
event with a 1,000- to 2,000-year mean return period and the lower level
Functional Evaluation Earthquake (FEE) with a 300-year mean return period.
Under a SEE, any damage to the bridge has to be readily repairable without
limiting traffic on the bridge. Under a more frequently occurring FEE the
damage would not require repair or cause any permanent offsets. As much as
possible, the Important Bridge criteria as required by Caltrans are met.
Performance criteria are set by limiting strain levels in concrete,
reinforcement and steel elements.

When designing for an FEE event a maximum concrete strain of
.004 was used for reinforced concrete elements. For a SEE event, the
stress-strain relationships proposed by Dr. J.B. Mander for confined concrete
were used. For the tower piles, this strain was limited to 75% of the ultimate
strains determined by Mander’s equations.

To achieve the performance goals under a FEE event, the
tension reinforcement strain was limited to .015. Similar limits were
established under the SEE event for ultimate strains as well as design level
strains. Limiting strains in the steel casing of pile shafts also was
established for pile compression as well as longitudinal and hoop tension.

Geomatrix Consultants, Oakland, carried out seismic ground
motion study for maximum credible earthquakes (MCE) on the governing faults.
Three design earthquakes were selected for the Carquinez Bridge.

Comparison of the response spectra of the MCE on the San
Andreas, Hayward and Franklin faults indicate that the spectral values for the
Franklin event are equal to or higher than the other two events over the entire
period ranging from 0 to 7.5 seconds. The response spectrum for this event was selected
as the design response spectrum. Three sets of multi-support rock motions based
upon the Franklin event were developed for the non-linear time-history analysis
of the bridge.

Rock motions were developed at each support location of the
bridge corresponding to the SEE. Near-source effects were incorporated in these
motions to account for directivity effects. Also included were spatial
variation in the motions over the length of the bridge.

 

Modeling for research

The seismic performance evaluation of the bridge made use of
a global inelastic dynamic computer model of the entire bridge. Seismic demands
obtained from the structural analysis were compared with the capacities of the
respective elements to ensure that their seismic performance is consistent with
the project design criteria.

The global analysis was supported by detailed component
analysis of the steel box girder deck, tower legs and tower piles. Local
component analysis helped provide insight into their performance resulting in
considerable design and detailing refinement.

 

Modeling: A global
model of the bridge was developed from the north to the south anchorage. The
model included the Crockett approach viaduct on the south side of the bridge,
which was designed by Caltrans. The model captured the interaction between the
approach viaduct and the suspension bridge.

The global model accurately represented the bridge geometry,
stiffness and mass and considered all important structural components including
pile foundations, anchorages, main towers, transition pier, cables, suspenders
and the orthotropic deck. To start with, a linear elastic model was used, but
this was refined during the analysis as geometric nonlinearity was introduced
in the cables and material nonlinearity was introduced in concrete elements
found to experience inelastic behavior.

In the time history analysis, which was based on the Newmark
Implicit Integration method, Rayleigh damping was used for the structural
elements with the damping matrix C=aM + bK. Here M is the mass matrix and K the
stiffness matrix. The mass proportional damping a and the stiffness
proportional damping b were determined for each element group by following an
iterative process.

Minor inelastic deformations were permitted at the base of
the tower leg and below the lower strut under extreme seismic loads.
Moment-curvature (M-F) relationships were developed and used at these locations
to model the material nonlinearity. Nonlinear M-F relationships also were used
to model the material behavior over the full length of the piles. The M-F
curves varied over the height of the piles with variation in reinforcement and
confinement provided by the cap, the shell and the rock.

Additional hydrodynamic mass and soil mass is added to the
pile elements through discrete lumped mass at the nodes. The hydrodynamic mass
was assumed to be equal to the volume of water displaced by the piles. The soil
mass tributary to the tower piles was assumed to be the volume of soil
contained within a “doughnut” of a radius equal to 1.5 times the
pile radius.

 

Analysis results:
The nonlinear time history analysis of the bridge was performed using the ADINA
program. Results obtained from this analysis were used to design the various
bridge components.

The tower struts, which are post-tensioned boxes, have been
designed to remain elastic during the SEE. The upper strut has been designed to
have a flexural capacity 1.2 times the tower leg plastic moment capacity. This
ensures inelastic deformations are limited to the legs. The lower strut has not
utilized the same criteria, since designing the strut for combined plastic
moments above and below the strut would be unrealistically conservative. The
lower strut has been designed for a nominal moment of 1.2 times the elastic
demands found from the global 
seismic analysis.

For the purpose of evaluation, the tower piles are assumed
to have three distinct sections:

1) Pile at the pile cap where the rebar cage extends into
the cap;

2) Pile below the cap and above the rock socket where it is confined
by a permanent casing 13/4 in. thick. Casing contributes to both lateral
confinement and flexural strength; and

3) Pile within the rock socket.

Pile flexural capacity was obtained from an M-F analysis for
varying axial loads. The capacity was based either on the limiting compressive
strains in the confined concrete or tensile strains in the reinforcement or the
permanent casing. In determining the capacity of the pile section with casing,
the shell thickness was reduced to account for corrosion.

The adequacy of the pile sections under seismic loads was
determined by plotting the curvature capacity of each section at varying loads.
Seismic curvature demands and the corresponding axial load on the pile at
various time steps was plotted on this same graph. If the capacity curve
enveloped the data points obtained from the seismic demands, the section design
was assumed to be adequate.

 

Local modeling:
Local structural models were developed to support seismic analysis and design.
Models of the tower bases, piles and pile caps were developed to verify design
capacities, section ductilities and evaluate compliance with the seismic
performance criteria. This advanced analysis work was performed by Anatech
Consulting Engineers, San Diego, and SCSolutions Inc., Santa Clara, Calif.

Local continuum analytical studies were performed for
lower/upper strut-to-column joint region and for the tower base/pile-cap/pile
system to verify this system’s seismic performance.

 

Ready to ride

The new Carquinez Bridge was designed for forces obtained
using state-of-the-art analysis methods. Rock motions at each support were
developed corresponding to the SEE incorporating the near-source effects. The
three-dimensional model of the bridge incorporated geometric nonlinearity in the
cables and material nonlinearity in the concrete elements which show inelastic
behavior. Soil-structure interaction was considered using foundation impedance
matrices.

Seismic studies demonstrated that the bridge design meets
the seismic design criteria adopted for the project. During a SEE event,
minimal damage is expected in the tower legs and tower piles, which could
result in cover concrete spalling and limited yielding of the reinforcement.
During a FEE event, no damage is expected to these bridge elements.

Suspension bridges are inherently flexible structures with
the suspended superstructure and tall towers resulting in long periods of
vibration. Thus, these bridges are well suited in areas of high seismicity. The
new Carquinez Bridge should therefore ride out the next big earthquake without
suffering major damage.

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