Case Study
Foundation Settlement and Tilt of
Millennium Tower in San Francisco, California
This work is made available under the terms of the Creative Commons Attribution 4.0 International license.
Jonathan P. Stewart, F.ASCE 1; Nathaniel Wagner, M.ASCE 2;
Debra Murphy, M.ASCE 3; Jeremy Butkovich, M.ASCE 4;
Micaela Largent, M.ASCE 5; Hamid Nouri, M.ASCE 6;
Hannah Curran, M.ASCE 7; Darcie Maffioli, M.ASCE 8;
and John A. Egan, M.ASCE 9
Abstract: The Millennium Tower is a 58-story reinforced concrete building that was constructed in San Francisco, California, between 2005
and 2009. The Tower is founded on an embedded pile-supported mat with pile tips bearing in dense marine deposits that overlie an overconsolidated marine clay layer known locally as Old Bay clay. This clay layer experienced stress increases from Tower self-weight and from
multiple episodes of dewatering between 2006 and 2018 at the Tower site and neighboring sites, including one that was sustained for 6 years.
Settlements of the Tower foundation have been measured since 2006, and lateral deflections of the Tower have been inferred and measured
since 2009. The data show that during multiple episodes of “loading” (from stress increase or dewatering), settlements initially accelerated
and then gradually slowed over time, as expected from consolidation principles. Similarly, lateral deflections (from foundation tilt)
(1) accelerated following foundation construction activities at adjacent sites (dewatering and excavation); and (2) oriented toward adjacent
excavations, which at various times occurred to the project south, north, east, and west of the Tower site. An objective of this paper is to
describe this case history, including the geotechnical site conditions and results of a monitoring program that tracked foundation settlements,
Tower tilt, groundwater levels at the Tower site, and ground inclinations over time. We also evaluate soil deformation mechanisms that likely
produced the movements. We find that settlement amounts and time variations are well captured by one-dimensional and three-dimensional
analyses of volume change in Old Bay clay and other foundation soils from primary consolidation and secondary compression, provided that
time variations of stress increase and groundwater level are accounted for. Three-dimensional analyses also capture time variations of lateral
deflections, which were caused by volume change and shear deformations in foundation soils, the latter having been affected by unloading
from adjacent excavations. DOI: 10.1061/JGGEFK.GTENG-10244. This work is made available under the terms of the Creative Commons
Attribution 4.0 International license, https://creativecommons.org/licenses/by/4.0/.
Author keywords: Millennium Tower, San Francisco, CA; Field monitoring; Foundation performance; Deep excavations; Groundwater
drawdown; Consolidation; Secondary compression.
1
Professor, Dept. of Civil and Environmental Engineering, Univ. of
California Los Angeles, Los Angeles, CA 90095 (corresponding author).
ORCID: https://orcid.org/0000-0003-3602-3629. Email: jstewart@seas.ucla
.edu
2
Project Engineer, Slate Geotechnical Consultants, 2927 Newbury St.,
Suite A, Berkeley, CA 94703. Email: nwagner@slategeotech.com
3
Principal Engineer, Slate Geotechnical Consultants, 2927 Newbury St.,
Suite A, Berkeley, CA 94703. Email: dmurphy@slategeotech.com
4
Senior Associate, Shannon & Wilson, Inc., 400 N 34th St. #100,
Seattle, WA 98103. Email: Jeremy.Butkovich@shanwil.com
5
Project Engineer, Slate Geotechnical Consultants, 2927 Newbury St.,
Suite A, Berkeley, CA 94703. Email: mlargent@slategeotech.com
6
Senior Engineer, Shannon & Wilson, Inc., 400 N 34th St. #100,
Seattle, WA 98103. Email: hamid.nouri@shanwil.com
7
Geosciences Engineer, Pacific Gas and Electric Company, 300
Lakeside Dr., Oakland, CA 94610. Email: hannah.curran@pge.com
8
Senior Engineer, Rockridge Geotechnical, 270 Grand Ave., Oakland,
CA 94610. ORCID: https://orcid.org/0000-0002-5006-3259. Email:
damaffioli@rockridgegeo.com
9
Senior Principal Engineer, Independent Consultant, 766 Brookside
Dr., Danville, CA 94526. ORCID: https://orcid.org/0000-0002-5591-1798.
Email: johnaegan13@gmail.com
Note. This manuscript was submitted on July 12, 2021; approved on
September 2, 2022; published online on March 27, 2023. Discussion period
open until August 27, 2023; separate discussions must be submitted for
individual papers. This paper is part of the Journal of Geotechnical
and Geoenvironmental Engineering, © ASCE, ISSN 1090-0241.
© ASCE
Introduction
The Millennium Tower (referred to herein as the “Tower” in the text
and as “MT” in figures) was constructed at the southeast corner of
Mission and Fremont Streets in San Francisco, California, between
2005 and 2009 (Fig. 1). The Tower has 58 floors (approximately
184 m tall) and one basement level. An adjoining midrise and podium (MP) were constructed concurrently with the Tower on the
same city block. Following Tower construction, major neighboring
developments were constructed from 2011 to 2017 to the project
south and west (Salesforce Transit Center, STC; known initially as
the Transbay Transit Center), to the project north (Salesforce East,
SFE), and to the project west (Salesforce Tower and Plaza, SFT
and SFP).
The Tower began to settle during construction, and by approximately March 2008, settlements had exceeded the original geotechnical report estimate. The rate of initial settlements slowed with time
following construction, and settlement estimates were updated in
February 2009 as part of a permitting process prior to Tower occupancy. Multiple episodes of accelerated foundation settlement and
tilt, which were recorded as part of a monitoring program, occurred
in subsequent years.
The settlement and tilt of the Tower attracted media attention
(e.g., CBS 2017), and eventually disputes arose between various
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Fig. 1. Site plan showing locations of Millennium Tower (MT), its adjacent podium (MP), and neighboring developments including the Salesforce
Transit Center (STC), Salesforce Tower and Plaza (SFT, SFP), and Salesforce East (SFE) tower. Also shown are locations of borings, piezometers, and
inclinometers discussed in this paper. True north and “project north” (aligned with Fremont Street) are indicated.
parties as to the cause of the foundation movements and what
(if any) mitigation measures should be undertaken. The data
and interpretations presented here were initially developed during
legal proceedings.
This case history has value as a data set of multiepoch recorded
foundation movements, several eras of which provide insights into
mechanisms by which dewatering and deep excavations can affect
adjacent structures and that can be used to validate various analysis
methods. The aims of this paper are to (1) objectively document the
foundation movement data and other data on groundwater levels
and ground inclination profiles so that the information can be accessed and used by geotechnical engineers, (2) interpret and analyze the data to illustrate the volumetric and shear soil deformation
mechanisms that caused the foundation movements, and (3) evaluate whether the observed movements could be predicted using onedimensional and three-dimensional analytical models. We do not
evaluate what could have been done differently during foundation
design, nor do we analyze hypotheticals in which certain elements
of the case history were not present. Moreover, we do not address
the design and construction of the foundation remediation for the
Tower, which is ongoing as of this writing.
Subsequent sections present site characteristics, structure characteristics, and results of field monitoring programs. We interpret the
collective data set and describe mechanisms of foundation movement suggested by the data. We describe one-dimensional analyses
that illustrate the extent to which volume change from consolidation
and secondary compression can explain observed settlement variations over time. We also present three-dimensional analyses that elucidate the impacts of combined volumetric and shear deformations
on settlements and lateral deflections of the Tower foundation.
Site Characteristics
Geotechnical site conditions (stratigraphy, soil properties) near the
Tower site have been investigated in a series of studies since 1967.
© ASCE
The most critical stratigraphic interval is an overconsolidated marine
clay locally known as Old Bay clay, for which extensive testing to
establish stress history, undrained shear strength, consolidation, secondary compression, and index properties has been performed. Soil
properties presented herein reflect the state of knowledge circa
2020, which is different from that at the time of foundation design.
Stratigraphy
Table 1 catalogs geotechnical investigations at the Tower and neighboring sites that were considered in developing models for layer
stratigraphy and soil properties. Fig. 1 shows locations of borings
and cone penetration test (CPT) soundings. A spreadsheet file in
the Supplemental Materials contains example boring logs (Slate
Geotechnical Consultants, Inc. 2019), and the “Data Availability”
section provides information on how to access additional borings,
CPT soundings, and other site data.
The stratigraphy at the Tower and neighboring sites is illustrated
in the cross-sections in Fig. 2. The north–south and east–west
directions indicated in the section are oriented relative to “projectnorth” (parallel to Fremont Street, Fig. 1). As shown in Fig. 2,
surficial soils at the site consist of heterogeneous fills ranging in
thickness from approximately 4.6 to 7.6 m. The fill was removed
during construction of the Tower and MP structures at the project
site. Underlying the fill is a soft to medium stiff marine clay deposit, known locally as recent Bay deposits or upper young Bay
mud (YBM), which ranges in thickness from 6.1 to 9.1 m (thicker
on the west side of the Tower site). Underlying the upper YBM is a
zone of dense clayey sand and sand with clay (labeled in Fig. 2 as
“Marine Sands”), which in turn is underlain by interbedded stiff to
very stiff sandy clay (labelled as “Lower Bay Mud”) and mediumdense to dense clayey sand and sand with clay (“Lower Marine
Sands” or “Colma Sands”) to depths of approximately 27.4–30.5 m
(elevation NAVD88 −22.9 to −25.9 m). This unit has historically
been commonly used as a foundation bearing layer in this portion
of San Francisco.
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Table 1. Catalog of geotechnical investigations at the Tower and neighboring sites
Location
No. of
borings
Max. depth
[m (ft)]
Lab tests
Citation
50 Beale
7
76.2 (250)
—
50 Fremont
5
45.7 (150)
350 Mission
2 (cone pene.
tests only)
15.8 (52)
301 Mission
7
67.1 (220)
M/D, G, AL, TXUU, ILC
STC
22
106.7 (350)
M/D, AL, G, ILC, CRS, DS, DSS,
TXICU, TXCK0 UC, TCUU, Sk
415 Mission
10
90.2 (296)
301 Mission
3
57.0 (187)
M/D, G, F, AL, ILC, CRS, TXICU,
TXICD, C
—
301 Mission
1
51.5 (169)
M/D, SG, AL, TXUU, TXICU, ILC
301 Mission
2
79.3 (260)
ILC, CRS, HC
Dames and Moore (1966), Boring logs for 50 Beale Street,
San Francisco, California, prepared for Bechtel Corporation
San Francisco Office Building by Skidmore, Owings & Merrill
(SOM), January 17 [From Treadwell & Rollo, Inc. (2012)].
Dames and Moore (1981), Geotechnical Investigation, Proposed
Five Fremont Center Project, San Francisco, California, prepared
for Metropolitan Bechtel Shorenstein, March 13 [From Treadwell &
Rollo, Inc. (2012)].
Treadwell & Rollo, Inc. (1997), Geotechnical Investigation,
350 Mission Street Building Seismic Strengthening, San Francisco,
California, Project No. 2152.01, July 3 [From Treadwell &
Rollo, Inc. (2012)].
Treadwell & Rollo, Inc. (2005), Revised Geotechnical Investigation
301 Mission Street, San Francisco, California, Project No. 3157.02,
January 13.
Arup North America, Ltd. (2010), Transbay Transit Center
Program: Transbay Transit Center, Contract No. 08-04-CMGC-000,
Volume 7A/B, February 26.
Arup North America, Ltd. (2013), Transbay Tower Final
Geotechnical Data Report, Job Number 229478-00, May 17.
Arup North America, Ltd. (2018), Memorandum, Salesforce Transit
Center, Supplemental Instrumentation Installation, March 20.
SAGE Engineers, Inc. (2018), Data Report for Geotechnical
Investigation and Piezometer Installation, 301 Mission Street, City
and County of San Francisco, California, November 27.
Slate Geotechnical Consultants, Inc. (2019), Geotechnical
Investigation & Instrumentation Installation - Plenum Area,
301 Mission Street, San Francisco, California, July 19.
M/D, AL, TXUU, DS/CD
—
Note: Laboratory test abbreviations: M/D = moisture/density; AL = Atterberg limits; G = gradation; F = fines percent; SG = specific gravity; Sk = slake
durability; C = corrosivity; HC = (falling head) hydraulic conductivity; ILC = incremental loading consolidation; CRS = constant rate of strain consolidation;
DS = direct shear; DSS = direct simple shear; DSCD = consolidated, drained direct shear, TXUU = unconfined, undrained triaxial compression; TXICU =
isotropically consolidated, undrained triaxial compression; TXICD = isotropically consolidated, drained triaxial compression; and TXCK0 UC =
K0 -anisotropically consolidated, undrained triaxial compression.
The Old Bay clay (OBC) unit underlies the lower marine/Colma
sands. The OBC is divided into upper and lower units (Arup North
America, Ltd. 2010), with the transition marked by a stiffer, sandier
crust occurring at depths between approximately 48.8 and 54.9 m
(elevation NAVD88 −44.2 and −50.3 m) depending on the location.
The thickness of the upper OBC unit ranges from 21.3 to 24.4 m,
being thickest (and encountered at the shallowest depth) on the west
and north sides of the structure (Fig. 2). The thickness of the lower
OBC unit ranges from 15.2 to 24.4 m depending on the location.
The lower OBC unit is also referred to as the Alameda formation.
Franciscan Complex bedrock underlies the lower OBC unit at depths
ranging from about 67.0 to 76.2 m (elevation NAVD88 −68.6 to
−71.6 m within the Tower footprint). The Franciscan Complex bedrock is a mélange matrix locally characterized by very soft, weathered shale and occasional blocks of serpentinite and sandstone.
Properties of Old Bay Clay Layers
The OBC layer within the site profile has been the subject of multiple programs of laboratory testing in the investigations listed in
Table 1. The soil specimens used in laboratory testing were derived
from samples of variable quality, having been obtained from splitbarrel driven samplers [standard penetration test (SPT), and larger
diameter versions known locally as California samplers] and various
pushed, thin-walled tube samplers (Shelby tubes, pitcher-barrel, various types of piston tubes). The investigations identified in the last
© ASCE
five rows of Table 1 were not available at the time of the Tower design, but these data are considered here because our aim is not to
interrogate the foundation design process but rather to document
and evaluate the case history given currently available information.
We compiled laboratory test results from the various investigations (Table 1), focusing on thin-walled pushed tube samples for
the case of consolidation and undrained strength properties. Consolidation properties compiled as a function of depth are (Fig. 3):
preconsolidation pressure (σp0 ), which is compared to preconstruction vertical effective stress in the figure; virgin and recompression
strain indices (CCε and Crε , respectively); secondary compression
strain index following an increment of virgin compression (Cαε );
and coefficient of consolidation following virgin or recompression
load increments (cv ). Consolidation properties of the layers overlying the OBC are not shown because the contributions to the overall settlement of the Tower from these layers are relatively small, as
shown subsequently.
As shown in Fig. 3, prior to construction, the upper OBC unit
was moderately overconsolidated [over-consolidation ratio (OCR)
∼1.8 to 2.3] in the upper ∼4 m, reducing to OCRs of 1.6 to 1.7 at
the base of this unit. This unit is also compressible (high compression indices) relative to lower OBC units. Secondary compression
properties of the upper OBC unit were investigated to characterize
response following primary consolidation (as in Fig. 3) and also
following stress reduction (Wagner et al. 2021). We investigated
whether OCR or compressibility parameters demonstrated a pattern
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Fig. 2. Interpreted cross-sections through Millennium Tower site and adjacent properties in project north–south and east–west directions (modified
from Arup North America, Ltd. 2010). Water levels shown in margins correspond to its limits within the lower marine (Colma) sands, as represented
in the timeline shown in Fig. 5.
Fig. 3. Profiles of consolidation properties of OBC (preconsolidation pressure, virgin and recompression strain indices, secondary compression strain
index, coefficient of consolidation). Lines indicate best estimates used in simulations. Coloration corresponds to layers in cross-section diagrams
(Fig. 2)—light blue is upper OBC and brownish green is lower OBC (Alameda); overlapping colors correspond to approximate depth variation
between layers.
© ASCE
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Undrained strengths derived from additional pocket penetrometer
and field torvane results are not shown in Fig. 4.
Structures
Fig. 5 shows a construction timeline for the Millennium Tower
and the three major nearby projects. The following sections
describe the Tower and pertinent foundation construction for adjacent projects. This information on features of the construction and
their timing is important for understanding linkages to foundation
performance.
This work is made available under the terms of the Creative Commons Attribution 4.0 International license.
Millennium Tower
Fig. 4. Profiles of stiffness and strength properties, as used in 3D
simulations. Best estimate profiles shown as lines. Coloration corresponds to layers in cross-section diagrams (Fig. 2).
of horizontal variability across the project site but did not identify
systematic trends.
Fig. 4 shows stiffness and strength properties of OBC layers as
reported in Arup North America, Ltd. (2010) (this report also includes such properties for other soil units). Stiffness is taken from
in situ shear wave velocity (V S ) from downhole measurements near
the Tower site. Indirect and direct strength data are compiled as SPT
blow count N (no energy or overburden correction applied) and laboratory and in situ undrained strength. The undrained strength data
points in Fig. 4 are from unconsolidated–undrained triaxial compression tests. In addition, consolidated–undrained test programs
that applied the SHANSEP approach (Ladd and Foote 1974) found
normalized undrained strength parameters (Su =σc0 ¼ S × OCRm ) of
S ¼ 0.23 and m ¼ 1.0 (Arup North America, Ltd. 2009). An undrained strength profile based on those parameters and an OCR profile derived from preconsolidation pressures are shown in Fig. 4.
The Millennium Tower is a 184-m-tall, 58-story, reinforced concrete
frame with a central core of shear walls. The substructure consists of
a one-level basement, 4.6 m deep, with a 3-m-thick foundation mat
over 945 driven reinforced concrete piles (DeSimone Consulting
Engineers, LLC 2005). The piles are driven into the lower marine/
Colma sand unit (Fig. 2). The piles have a 35.6-cm square cross
section and center-to-center spacing ranging from 1.07 to 1.42 m.
Immediately adjacent to the Tower is an 11-story (38-m-tall)
midrise and three-story podium structure. As shown in Fig. 2, the
MP substructure consists of a four-level basement, 16.7 m deep,
supported on a shallow mat foundation bearing on the upper marine
sand layer. The Tower and MP are separated by a soil–cement mixture cutoff wall containing embedded steel H-beam soldier piles,
installed during construction of the MP. Fig. 6 shows a foundation
plan for the Tower.
Tower construction commenced with indicator pile installation
in late 2005, which was followed in early 2006 by installation of
production piles (Treadwell & Rollo, Inc. 2006), installation of the
MP cutoff wall, and excavation dewatering for the Tower and MP
sites (Fig. 5). As shown subsequently, this dewatering temporarily
lowered the groundwater at the site by 9–10 m. The Tower mat
was poured at the end of June 2006. Construction of the Tower
Fig. 5. Construction timelines at Millennium Tower site and adjacent properties.
© ASCE
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Fig. 6. Millennium Tower foundation plan showing pile layout and configuration of lateral force resisting elements. (Adapted from Hamburger
et al. 2021. Reprinted with permission, Structure, June 2021.)
superstructure started in August 2006, followed by excavation for
the MP starting in September 2006. Tiedown anchors were installed
through the foundation mat on the west side of the MP to counteract
buoyancy effects following groundwater rebound. Dewatering ended
in February 2008, and superstructure construction finished in late
2008. The Tower was opened for residents in April 2009.
Adjacent Construction
Following construction of the Millennium Tower, three projects were
completed at adjacent sites (Fig. 1). The Transbay Transit Center
(now Salesforce Transit Center) is located to the project south and
© ASCE
west and was constructed from 2010 to 2018. A structure at 350
Mission (now Salesforce East) is located to the north and was constructed from 2013 to 2015. The Salesforce Tower and Plaza are
located to the west and were constructed from 2015 to 2018.
For the purpose of understanding Tower foundation performance, the most pertinent aspects of adjacent construction are related to dewatering and excavation, which are highlighted in the
Fig. 5 timeline. Major elements of substructure construction include installation of cutoff walls along the project perimeter that
are intended to have limited conductivity and thereby maintain
groundwater levels beyond the project perimeter as dewatering
occurs within the site during excavation and substructure or
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superstructure construction. Dewatering is normally terminated
immediately following substructure construction, although in the
case of the STC, the dewatering continued beyond the period of substructure construction, as discussed further subsequently.
Transit Center
The STC project extends over four blocks, and the construction was
subdivided into four zones, as shown in Fig. 1, with Zone 3 being
immediately true south of the Tower (i.e., project southwest) and
Zone 4 being adjacent to and true southeast (i.e., project south) of
the Tower. The information on the STC timeline in Fig. 5 is from
Transbay Joint Powers Authority (TJPA 2019). Demolition of the
former Transbay Terminal started in mid-2010, and cement deep
soil mixing (CDSM) cutoff walls were installed throughout 2011
and into early 2012 (including pretrenching, timber pile removal,
and shallow dewatering). A prototype CDSM wall was installed at
the west end of the site to test the construction procedure and resulting material properties (Kluzniak et al. 2010). Dewatering began in March 2012 for Zones 1–2 and in May 2013 for Zones 3–4.
Excavation to approximately 13.5 m depth (Fig. 2) started in Zone
1 in early 2012 and moved east toward Zone 4 through early 2014;
excavation in Zone 4 (adjacent to the Tower) started in early 2013.
Prior to the excavation in Zone 4, a subsurface buttress was installed adjacent to the Tower from mid-2011 to early 2013 (Arup
North America, Ltd. 2009; TJPA 2019), which was intended to
limit Tower foundation movements related to STC construction. The
buttress consists of 182 drilled shafts 2.1 m in diameter extending
into the Franciscan Complex. Substructure construction was completed by early 2014 for Zones 1–2 and late 2014 for Zones 3–4,
which was followed by superstructure construction. Dewatering was
not concluded upon the completion of substructure construction, instead extending to mid-2018 when the STC was opened.
350 Mission
The SFE project developed a 129-m-tall, 30-story building. The
substructure consists of a four-level basement, 15 m deep, and a
2.5-m-thick foundation mat that bears on the upper marine sand
layer. Demolition of the prior structure occurred in early 2013,
followed shortly thereafter by cutoff wall construction (similar in
configuration to that for MP) and dewatering and excavation that
essentially coincided with dewatering of Zones 3–4 and excavation
of Zone 4 at the STC site. Substructure construction began in late
2013 and was followed by superstructure construction that began in
early 2014.
Salesforce Tower and Plaza
The project has two major components: a 326-m-tall, 61-story
tower supported on a 15.2-m-deep, four-level basement (Salesforce
Tower), and a ground-level Plaza that consists of a 10.6-m-deep,
two-level subsurface garage extension (Salesforce Plaza). As shown
in Fig. 1, SFP is located between SFT and the Millennium Tower.
The SFT foundation consists of a barrette foundation-supported mat,
with the barrettes extending into bedrock. The SFP foundation is a
1.5-m-thick foundation mat. For the SFT, cutoff wall construction,
dewatering, and excavation started at the end of 2014 and continued
through 2015. The foundation mat and substructure construction occurred from late 2015 through early 2016, with superstructure construction occurring through late 2017. Excavation and dewatering
for the SFP occurred in early 2017, and basement construction was
completed in 2017.
Performance Monitoring
Millennium Tower foundation settlement at limited benchmarks
has been monitored since Tower construction began in 2006. Subsequently, as construction occurred at neighboring properties, numerous additional settlement markers and instruments were installed at
and around the Tower, including piezometers, inclinometers, extensometers, survey prisms, and tiltmeters. In this section, we present
results from representative piezometers, inclinometers, settlement
markers, and survey prisms to describe groundwater variations,
foundation movements, and soil deformations at the Tower site for
time intervals with available data from 2006 to 2020.
Groundwater
The preconstruction groundwater table in the vicinity of the Tower
occurred at a depth of approximately 3 m (Treadwell & Rollo, Inc.
2005). As shown in Fig. 7, starting with the construction of the
Tower in 2006 and continuing through 2018, the water table was
lowered by a series of dewatering programs at and adjacent to the
Tower site. The significance of dewatering for the Tower foundation
performance is that it reduces buoyancy of surficial soils (fill and
young Bay mud) and of the Tower, thus increasing effective stresses
Fig. 7. Variation of groundwater depth representing piezometric head in the upper marine sand and lower marine/Colma sand with time at the Tower
site as measured from wells (2007) and various piezometers (mainly 2011–2020). Also shown for the time period of 2006–2011 is piezometric
head in dune sand at the west end of STC site during Tower construction. Colored circles indicate data obtained manually in the respective layers.
The “Baseline Analysis” depths represent an interpreted spatial average within the Tower foundation footprint.
© ASCE
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on deeper layers, including the upper OBCs. Whereas piezometers
have been installed in various soil units, Fig. 7 shows groundwater
depths associated with measurements in the upper/lower marine
sand and Colma sand units that directly overlie the upper OBC.
Dewatering activities during substructure construction of the
Tower and MP were tracked by water table depth measurements in
a series of standpipes, including nine wells used for dewatering in
the MP excavation and two wells on the Tower side of the cutoff
wall separating the two excavations. Measuring ground water depths
from the same wells used for dewatering overestimates drawdown
relative to the expected value within excavation areas due to local
cones of depression at the wells from pumping. Nonetheless, the
available data indicate that the tips of “cones” were at about 23–26 m
depth within the MP cutoff walls, whereas depths measured in the
two wells on the Tower side of the cutoff wall separating the excavations were approximately 12–13 m (the difference is likely influenced by limited conductivity through the cutoff walls). These
measurements establish the baseline analysis water table depth
shown in Fig. 7 for the period March 2007–February 2008.
Piezometers west of the Tower site outside of cutoff walls recorded
1–1.5 m of drawdown and tracked subsequent recovery, which was
essentially complete by early 2011 (Fig. 7).
Piezometers were installed near the Tower in 2011 to monitor
the effects of construction at the STC site, at the end of 2014 to
monitor the effects of construction at the SFT and SFP site (Arup
North America, Ltd. 2015–2017), and from 2016 to 2019 (Arup
North America, Ltd. 2018; SAGE Engineers, Inc. 2018; Slate
Geotechnical Consultants, Inc. 2019). The data from these piezometers are not sufficient in their spatial coverage to establish the
three-dimensional (3D) groundwater surface across the Tower site,
which is expected to be nonuniform. However, they provide valuable representations of water table elevation at multiple points that
allow patterns in time to be identified. As shown in Fig. 7, following some initial fluctuations in these data from mid-2011 to March
2012, a series of reductions in piezometric head in the upper and
lower marine sand and Colma sand units occurred as follows:
• April–May 2012: ∼1.5 m drawdown that coincided with the
start of STC Zone 1–2 dewatering.
• Throughout 2013: ∼1.3 m drawdown that coincided with the
start of STC Zone 3–4 dewatering and SFE dewatering.
• November 2014–February 2015: ∼1.5 m drawdown that
coincided with start of SFT dewatering.
• Start of 2017: ∼0.6 m drawdown that coincided with start of
SFP dewatering.
As shown in Fig. 5, dewatering at the STC site was sustained
through 2018, whereas other dewatering programs were relatively
brief. The effects of the sequenced dewatering were a cumulative
drawdown of approximately 5 m by early 2017 (i.e., drawdown to a
depth of approximately 8 m), which began to recover in mid-2017
following the completion of SFP dewatering. As of mid-2020,
groundwater had recovered to approximately 4.9 m depth (1.9 m
of residual drawdown).
In addition to the piezometers in the upper and lower marine
sand and Colma sand units used to monitor groundwater drawdown,
several arrays of piezometers were installed in the upper OBC very
close to the building perimeter (Arup North America, Ltd. 2018;
SAGE Engineers, Inc. 2018) and within the building envelope (Slate
Geotechnical Consultants, Inc. 2019) to investigate the possible
presence of excess pore pressures (i.e., water pressures exceeding
hydrostatic) that if present would be associated with ongoing consolidation at the time the measurements were made. Data from such
piezometers are available from May 2018 and May 2019 onward,
respectively. Fig. 8 shows profiles of water pressures measured in
April 2019 and May 2020. Although there are some piezometers
that indicate water pressures that exceed hydrostatic (e.g., Arup
piezometers at 37 and 42 m), the preponderance of evidence suggests that excess water pressures are essentially nil.
We interpret the multiepoch groundwater lowerings that coincide with dewatering at adjacent sites as an indication that there
was “communication” between water levels at the Tower site and
those within cutoff walls at adjacent sites. Such effects could be
caused by potential leakage through cutoff walls or flow beneath
the walls.
The interpreted groundwater level for use in the onedimensional (1D) analysis (Fig. 7) was established as described
previously for the period 2006–2010 and was interpreted from piezometric data for the period since 2011. The interpretation considered the proximity of piezometers to the Tower site and as such is
not a simple average of the measured water table depths. The interpreted groundwater levels in the upper and lower marine sand
and Colma sand units were considered representative for use as
a boundary condition in the 1D consolidation analysis. Minute
weekly or monthly fluctuations were not considered in the interpretation; instead, piecewise linear approximations were used to
smoothly transition between major drawdown episodes. The interpretation was done “by eye” following the general trends of
the data.
As a result of the excess water pressures at the site being essentially nil (Fig. 8), it appears that primary consolidation of the OBCs
had effectively concluded by mid-2020, and possibly earlier.
Foundation Settlement
Foundation settlement has been recorded since September 2006
using three monitoring programs: (1) September 2006–February
2009: a monument mounted on the first floor of the Tower core wall
Fig. 8. Profile of water pressures in Old Bay clays immediately adjacent to and directly beneath Millennium Tower foundation mat, as measured in
April 2019 and May 2020.
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Fig. 9. Plan view of Tower foundation mat showing settlement marker
locations. Markers for which settlements versus time are plotted in
Fig. 10 are shown in bold. No settlement marker was labeled SM-12.
was surveyed by Martin M. Ron Associates, Inc. at approximately
monthly intervals, (2) May 2009–present: 31 settlement markers
(SM-XX, where XX is the settlement marker identification number)
were installed on the Tower mat by Arup North America, Ltd.
and surveyed to record vertical position at approximately monthly
to bimonthly intervals through January 2017, weekly intervals
March 2017–February 2019, and roughly 1.5-month intervals since
February 2019, and (3) January 2017–present: the SM markers were
supplemented with an additional 30 settlement markers (LE-XX,
where XX is the settlement marker identification number) installed
on the Tower mat by Langan Engineering. The locations of the markers are shown in Fig. 9.
Fig. 10 shows the variation of settlement with time from the
monitoring program. The single-location data set through 2008 was
merged with the survey monument data by linking the MM Ron
data from the wall monument to displacement at SM-27, which
is the closest settlement marker. The merging procedure was verified by comparing settlements for overlapping time intervals. The
initial (as of April 2009) settlement at every other marker was calculated relative to SM-27 using elevation differences established
using a manometer survey. The LE data set is based on the same
elevation datum as the SM data set, so a merging procedure was not
required. As shown in Fig. 10, there are vertical shifts in the data for
different settlement markers, but the pattern with time is common to
each of them.
We represent the position of the foundation mat at a particular
time with a plane fit to the 3D survey data from the monuments
using bilinear least squares regression. By doing so, the vertical
position (settlement) and angles relative to horizontal in two reference directions (tilts) can be established at each time that settlement
markers were surveyed. Although the mat is not perfectly planar
(it has a low point in the northwest quadrant), analysis of residuals
between actual position and planar fit indicate a misfit standard
deviation increasing only slightly (from approximately 1.25 to
1.50 cm) between the start of monitoring in 2009 and 2020 and
a reasonably consistent contour map of residuals with time. In establishing the planar fit, we exclude settlement markers that were
inaccessible at some of the survey times (SM-13), those located
on walls instead of the foundation mat (SM-15 and 16), and those
located on a cantilevered and relatively thin (0.3 m thickness) portion of the foundation mat (noted in Fig. 9) that is not supported by
piles (SM-3, 4, 5, 6, and 7).
The mean settlement from the planar fit is marked in Fig. 10,
and falls near the middle of the range from individual settlement
markers. The planar fit to the settlement marker data also establishes foundation slab tilt that can be measured in any horizontal
direction. We denote tilt angles in the reference north–south and
east–west directions (parallel to Fremont and Mission Streets,
Fig. 10. Settlement of Millennium Tower foundation mat versus time at various settlement markers and from mean of planar fit. Directions in legend
are “Project” directions (Fig. 1). The planar mean settlement is the central value that provides the most representative settlement for performance
assessment and comparison to simulations.
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respectively) as θns and θew . We interpret the time-dependent tilt
established by this data in the next subsection.
Our interpretations of notable features of the settlement-time
data in Fig. 10 are as follows:
• September 2006–February 2008: During this period of local
dewatering and Tower construction, settlements developed rapidly but decelerated toward the end of the construction interval.
• February 2008–April 2011: During this period following Tower
and MP construction but prior to other major construction activities adjacent to the Tower, settlements slowed with time.
• April 2011–April 2012: During this period, construction activities commenced at the STC site, including CDSM wall installation (April 2011) and Zone 1–2 dewatering (April 2012).
Settlement rates moderately increased.
• Early 2013, December 2014, February 2017: During each of
these three periods when dewatering and excavation commenced
adjacent to the Tower, settlements initially accelerated and then
gradually slowed with time.
Tower Tilt
We define tilt of the Millennium Tower as the vertical angle of the
vertical faces of the building. Tilt is denoted Θns and Θew for the
project north–south and east–west directions, respectively. As shown
in Fig. 11, tilt is directly related to horizontal deflection (Δ) as
Δ¼z×Θ
ð1Þ
where z = vertical distance from the base slab to the position where
horizontal deflection is measured. Eq. (1) applies if building deflections from tilt do not appreciably distort structural members, which
has been confirmed by structural analyses (Hamburger et al. 2021).
Three sources of data have been used to estimate the lateral deflection or tilt and their variations with time:
1. External survey data (January 2017 to present): Consist of total
station surveys of prisms on east and west sides of the Tower at
the top of the 2nd, 5th, 14th, 20th, and 40th floors and the top of
the façade (heights of 5.2, 14.6, 40.8, 58.5, 121.0, and 180.4 m
above street level). For each prism, location is provided (latitude,
longitude, elevation) with high precision (uncertainty < 0.1 cm).
2. Interferometric synthetic-aperture radar (InSAR) data (May 2009
to March 2017): Provide geodetic coordinates of a point on the
roof of the Tower as a function of time. Referred to herein as
“satellite data,” this information was provided by 3v Geomatics
as horizontal (easting and northing) and vertical movement relative to an initial reading in May 2009 (InSAR would not be
Fig. 11. Schematic illustration of tilt angle relationship to horizontal
deflections of survey prisms.
© ASCE
useful during noisy construction periods preceding this date).
As a consequence of the satellite flight path, north–south movements are more accurate (uncertainty of 2.8 cm) than the east–
west movements (5.8 cm). As shown subsequently, these uncertainties introduce transient data scatter but do not obscure longterm data trends (particularly when movements are much larger
than the uncertainties).
3. Horizontally gridded foundation mat settlement data (May 2009
to present): Data are described in the previous subsection. Planar
fits of data were used to establish foundation tilt angles relative
to horizontal, θns , and θew at each measurement date.
The Tower vertical tilt angles Θns and Θew , could potentially
differ from the foundation horizontal tilt angles θns and θew . A departure in these angles would occur if the Tower mat foundation
underwent tilt simultaneously with settlement during construction,
in which case incremental releveling of the structure during construction may have occurred.
To present results on directional lateral deflection (or tilt) versus
time, we require an estimate of tilt at the onset of measurements,
which is May 2009 from the InSAR and settlement data. We do not
consider foundation slab angles θns and θew suitable for this purpose
(due to the potential for correction of tilt during construction) and
instead adopt the following approach:
1. Begin with tilt angles Θns and Θew from January 2017 provided
by external survey data.
2. For times prior to January 2017, estimate changes in tilt angles
dΘns and dΘew using InSAR and mat settlement data, as follows:
• Changes in time of InSAR-based directional horizontal deflections (dΔ) are converted to directional changes in tilt as
dΘ ¼ dΔ=z.
• Changes in time of directional mat tilt are taken as equivalent
to changes in directional Tower tilt, that is, dθ ¼ dΘ.
3. Compute tilt for time t as
ΘðtÞ ¼ Θðtr Þ − dΘðt − tr Þ
ð2Þ
where tr = reference time with known tilt (January 2017).
4. Compute the deflection of the highest survey prism using Eq. (1).
The vertical distance (z) in this case is 185.0 m above the mat
level.
Using the previous procedure, the variations of roof position in
the horizontal plane over time since May 2009 are shown in Fig. 12.
Fig. 13 shows the roof deflection trajectory in the horizontal plane
since May 2009, with line coloration indicating different construction intervals, as shown in Fig. 12. Separate results are shown for
survey, InSAR, and settlement monument data. The estimated deflections in May 2009 (at the onset of measurements from InSAR
and basement monuments) are 1.3–7.5 cm to the east and 10–12 cm
to the north.
Our interpretations of notable features of the deflection-time
data in Figs. 12 and 13 are as follows:
• May 2009–April 2011: During this period following Tower/MP
construction, the planar estimate data (from foundation settlements) indicate lateral movements progressed toward the project
north and west at a relatively slow pace. InSAR data are relatively sparse in time during this interval.
• April 2011–April 2012: During this period in which construction
activities commenced at the STC site, deflections in the project
north–south direction were redirected toward the south, and deflections toward project west modestly accelerated.
• 2013–2014: Following the onset of dewatering and excavation
in Zones 3–4 of the STC (south and west of the Tower) and SFE
(north of the Tower) in early 2013, project north–south deflections reoriented toward the north and project west deflections
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Fig. 12. Time variations of horizontal deflections at the location of the top prism on Millennium Tower (185 m above foundation mat) since May
2009 for (a) project north–south direction; and (b) project east–west direction.
Fig. 13. Roof deflection trajectory in horizontal plane from exterior surveys, satellite data, and planar fits of foundation settlement and tilt for the time
interval May 2009 to July 2020. Colors indicate different construction time intervals (Fig. 12). The vertical and horizontal axes show deflections in the
project north–south and east–west directions, respectively, which are defined in Fig. 1.
substantially accelerated. Deflections in both directions slowed
throughout most of 2014.
• December 2014, February 2017: Following the onset of dewatering and excavation at SFT and SFP (west of the Tower),
deflections accelerated toward the west, whereas north–south
deflections did not change. These westerly deflections have
slowed since 2018.
Potential mechanisms causing these deflections are discussed in
the “Performance Assessment” section subsequently.
Inclinometers
Inclinometers were installed near the Tower, mainly on the south
side (to track ground deformations related to STC construction) and
west side (to track ground deformation related to SFT and SFP construction). Locations of inclinometers are shown in Fig. 1, and they
were installed between 2011 and 2016.
© ASCE
Some difficulties were encountered in analysis of inclinometer
data south of the Tower; we did not have access to raw data, and the
available processed data had been adjusted in a manner that cumulative horizontal displacements (relative to the base of the inclinometer) could not be evaluated. These issues were not encountered in
data collected west of the Tower, the results of which are shown in
Fig. 14. The figure shows cumulative horizontal displacements relative to the base of the profile at different times.
The inclinometer data in Fig. 14 track horizontal ground deformations in the area between the Tower and SFT and SFP on the
west side of Fremont Street. Displacements are referenced to baseline readings at the onset of construction in January 2015. Small
displacements toward the east had occurred by July 2015, perhaps
as a result of tiebacks or bracing installed during SFT excavation.
The December 2016 profile shows conditions prior to SFP excavation, which include modest westward deflections within the depth
range of the SFT excavation and at greater depths. Following SFP
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deformation. Volume change is expected from effective stressdependent void ratio change (primary consolidation) triggered by
construction activities that induce effective stress increase (as pore
pressures dissipate) in the upper OBC. In addition, time-dependent
void ratio change (volumetric creep) coincident with and following
effective stress-dependent void ratio change is also expected.
Although creep is often modeled as secondary compression that follows primary consolidation, creep actually occurs simultaneously
with effective stress-dependent volumetric change (e.g., Taylor and
Merchant 1940; Bjerrum 1967; Mesri and Choi 1985; Kutter and
Sathialingam 1992; Terzaghi et al. 1996; Brandenberg 2017). Further
discussion is provided in Section S1 in the Supplemental Materials.
Shear deformations in upper OBC and other foundation soils are expected from building self-weight and neighboring excavations and
would contribute to Tower foundation movements.
The purpose of this section is to interpret the performance
monitoring data to evaluate the soil deformation mechanisms that
caused the foundation movements. Although they are informed by
first principles pertaining to volumetric and shear deformations,
these interpretations are independent of simulations. Simulations
are useful, but it is always possible to question simulation results
because of sensitivity to uncertain input parameters and boundary
conditions.
Movement Patterns Preceding Adjacent Construction
Fig. 14. Inclinometer data west of Millennium Tower prior to and
during SFP construction (between January 2015 and August 2017).
There is no reading from I-07 from December 2016.
excavation, westward deflections increased markedly over the depth
range of the excavation. Potential mechanisms producing these deflections are described in the next section.
Performance Assessment
Performance monitoring data, in combination with the timeline
of construction at and near the Tower site, can be interpreted to
provide insight into the soil deformation mechanisms that caused
the Tower’s foundation settlement and tilt since 2006. Anticipated
soil deformation mechanisms include volume change and shear
Settlement
Fig. 15 shows (with a dashed line) foundation settlements with respect to log time from the MM Ron data and SM-27 for the time
period September 2006 to March 2012. This time period precedes
large-scale dewatering from adjacent construction. The data present
a classical shape for a process of primary consolidation followed
by secondary compression (e.g., Fig. 9.16 of Holtz et al. 2011),
with an apparent transition to secondary compression from primary
consolidation at tp ≈ 530 days (March 2008). The deceleration of
settlement at the apparent tp is influenced both by groundwater recovery (reducing effective stress in the upper OBC) and pore pressure dissipation. The slope of the settlement–log time curve beyond
tp is approximately 16 cm per log cycle of time (range considering
alternate start or end times for fitting is 15–18 cm= log cycle time).
If interpreted as possible secondary compression, this rate would
be equivalent to H × Cαε , where H is the soil layer thickness
Fig. 15. Foundation settlement with respect to log time from the MM Ron data and SM-27. Dashed line before March 2012, solid line after March
2012. The product HCαε is the slope of the settlement–log time data as indicated in the figure.
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0 ), sum of σ 0 and 2009–2010 total stress increase, and preconstruction preconsolidation
Fig. 16. Stress profiles for preconstruction condition (σv0
v0
0
pressures (σp ). Stress increases are shown for Method 1 (equivalent raft and 2:1 method) and Method 2 (3D simulations), which are described in the
next section. The average pile tip depth is shown for reference.
contributing to secondary compression, and Cαε is the secondary
compression strain index in effect following primary consolidation.
Secondary compression rates depend on whether the preceding
primary consolidation was recompression of overconsolidated soil
or virgin compression. Fig. 16 shows the effective stress profile
data from Fig. 3 (left frame), three interpreted σp0 profiles, and two
estimates of the stress increase from net building weight and dewatering following initial loading (2009–2010) added to the effective
stresses from soil self-weight (Methods 1 and 2 used to develop
these stress increases, including the effects of the piles, are described in the next section). These results show that the upper OBC
was likely brought into virgin compression over the approximate
depth range of 33 to ∼40 to 45 m. If H is estimated using the full
thickness of the upper OBC (∼23 m), the Cαε implied by the data is
0.0071. Alternatively, if H is taken as the thickness of the normally
consolidated depth interval (∼7 to 12 m), then the implied Cαε is
0.013–0.022. The range of Cαε for the upper OBC unit from
laboratory tests in which secondary compression followed virgin
compression is 0.006 to 0.016, with an average of 0.010 (Fig. 3).
The implied Cαε from field performance data only modestly exceeds
that from laboratory tests, suggesting that whereas some primary
consolidation likely remained during this period (approximately
January 2009 to March 2012), secondary compression was a strong
contributor to the settlement. This hypothesis is evaluated further in
the “Simulations of Foundation Movement” section.
The interpreted σp0 profiles in Fig. 16 were used in simulations
of consolidation settlement, discussed in the next section. The position of the baseline profile was set: (1) to capture the trend of the
data, with emphasis given to samples taken at the Tower site prior to
construction (highlighted in Fig. 16), (2) to include a relatively
more overconsolidated crust at the top of the upper OBC from approximately 27–32 m, the presence of which is well established
from consolidation and CPT data (from reports cited in Table 1),
and (3) to produce a reasonable match between simulated and observed settlements. Because the interpreted profile was established
in consideration of each of these criteria, it does not necessarily pass
through the mean of the full body of σp0 data at all depth ranges.
Most notably, the baseline profile does not appear to go through
the mean of the σp0 data for the 30–39 m depth interval, although
it is broadly consistent with the highlighted points from the Tower
site. The high and low profiles bracket the baseline σp0 profile and
are used to investigate sensitivity of simulated settlements to σp0
© ASCE
uncertainties, as discussed further in “Simulations of Foundation
Movement” section.
Tower Tilt
Figs. 12 and 13 show that Tower movements in May 2009 were
approximately 10–12 cm to the north and about 1.3–7.5 cm to the
east. Subsequently, tilt continued toward the north and reversed
direction toward the west at a relatively steady rate in the 2010–
2011 time frame. The northward tilt is likely a consequence of nonuniform stratigraphy (shallower depth to OBC layer to the north, as
seen in Fig. 2). Potential causes for the changes in east–west direction tilt are more nuanced, and the subject of our interpretation here
is whether the observed directions of movement could be reasonably expected.
The initial (i.e., May 2009) tilt toward the east is consistent with
expectations. As shown in Fig. 17(a), groundwater drawdown
Fig. 17. Schematic illustrations of (a) nonuniform groundwater lowering adjacent to dewatered excavation; and (b) stress increase and
decrease on upper OBC unit from Tower and podium.
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related to dewatering at the adjacent podium was likely nonuniform
beneath the Tower footprint, with the lowest levels closest to the
east cutoff wall (the lowering near the wall would occur due to leakage through the wall or underseepage). Although the nonuniform
groundwater drawdown during this time interval cannot be confirmed explicitly, the spatial distribution is well informed by similar repeated observations during other time intervals, as described
in the “Groundwater” section. The larger the drawdown, the greater
the stress increase in the upper OBC driving the consolidation
process—hence, more volume change would be anticipated beneath the east side of the structure than the west, producing tilt.
Moreover, shear deformations associated with the MP excavation
would be expected to produce tilt.
Following the completion of MP construction and groundwater
recovery, the mechanisms producing eastward tilt ceased. At this
stage, the main driver of tilt is expected to be nonuniform stress
increase on the upper OBC from the combination of Tower weight
(producing loading) and the overcompensated MP (the difference
between excavated soil weight and MP structure weight amounts to
an average unloading of approximately 140 kPa within the podium
footprint). As shown in Fig. 17(b), this nonuniform stress change
produced a greater stress increase on the west side of the Tower,
which would be expected to produce gradual westward tilt tracking
the foundation settlement, as observed.
Movement Patterns since Adjacent Construction
Settlement
The solid line in Fig. 15 (after March 2012) shows foundation settlements with respect to log time following the onset of significant
dewatering activities at adjacent sites. There is clear acceleration of
the preceding settlement rate, which coincides with groundwater
lowering in 2012 (Fig. 7). We interpret this acceleration of settlement
as a renewal of primary consolidation associated with dewatering.
As shown in Fig. 7, multiple episodes of dewatering affected
the water table at the Tower site between 2012 and 2017, each of
which would be expected to produce additional primary consolidation by increasing the effective stress below the Tower. Among
those episodes, the STC dewatering was likely the most consequential, because it was maintained for 6 years. Fig. 18(a) shows
the settlement since April 2012 with the log time scale reset to
correspond with a new stage of primary consolidation. As with the
initial settlement–log time data starting in 2006 (Fig. 15), the data
since 2012 exhibit the characteristic shape expected from primary
consolidation that gradually transitions toward secondary compression. Although the shapes are similar, the time span of primary
consolidation is much longer in Fig. 18(a) (tp ∼ 1,900 days) than
in Fig. 15 (tp ∼ 530 days). This difference results from different
time intervals over which the loading that induced primary consolidation was applied. Initially, the loading was applied over a relatively short time interval (2006–2007) from Tower construction and
MP dewatering, whereas the later stages were applied from multiple
stages of dewatering at adjacent sites between 2012 and 2017.
Fig. 18(b) expands the view for a recent time interval (1,000 to
3,000 days since April 2012) along with a line fit to the final data
stage with slope H × Cαε ¼ 11.5 ðrange 10–12Þ cm= log cycle time.
This slope, which characterizes the settlement rate since late 2017, is
smaller than that found for 2009–2011 and is relatively consistent
with laboratory-based estimates of Cαε . This suggests that primary
consolidation had effectively ended by that time, which is also
consistent with lack of excess water pressures from piezometers (Fig. 8).
We recognize that resetting the start time for consolidation to
April 2012 affects the logarithmic slope calculations (Brandenberg
2017). An alternative that was considered was to reject the 2012
reset by maintaining the 2006 start time, which results in H ×
Cαε ¼ 20 ðrange 10–30Þ cm= log cycle time for the same time interval considered in Fig. 18(b). This rate exceeds considerably that
in Fig. 15, when consolidation may not have been complete, which
would require incomplete consolidation in the period since late
2017. Because the pore pressure data indicate that consolidation
is complete during this period (Fig. 8), we conclude that effective
reset in 2012 is the more reasonable interpretation.
Tower Tilt
Figs. 12 and 13 show that Tower movements toward the west accelerated at several times, including 2012, 2013, 2015, and 2017.
At each of these times, dewatering and excavations occurred west
of the Tower site (STC, SFT, SFP). In the north–south direction,
following very gradual northward tilt pre-2011, tilt abruptly changed
Fig. 18. (a) Foundation settlement with respect to log time from the MM Ron data and SM-27 since March 2012. Horizontal scale assumes reset of
time for the consolidation process in March 2012, which coincides with the commencement of dewatering in STC Zone 1 and 2; and (b) relatively
detailed settlement versus time data as the consolidation process transitions to secondary compression following the end of dewatering.
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direction toward the south in 2011 (commencement of STC construction), toward the north in 2013 (SFE and STC construction),
and effectively did not change from mid-2014 onward (during SFT
and SFP construction).
When the post-2011 and pre-2011 tilt observations are viewed
in aggregate, a pattern emerges in which the Tower consistently
tilts toward dewatered excavations at adjacent sites. This response
can be explained based on nonuniform volume change [Fig. 17(a)]
and shear deformation effects. Simulations are required to provide
insight into the relative significance of these effects, which is the
subject of the next section.
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Simulations of Foundation Movement
The emphasis of this paper to this point has been on the presentation
and interpretation of objective data on the Millennium Tower foundation movement case history. Modeling of the foundation performance is taken up in this section. We acknowledge that these
analyses are necessarily somewhat subjective and other modelers
might make different assumptions and produce different findings.
Our main objectives with the simulations presented here are to:
(1) evaluate the degree to which methods used in practice can capture measured settlements and their variation with time; and (2) provide insight into the relative significance of volume change and
shear deformations on foundation settlements. The second objective is investigated by evaluating differences between foundation
movement predictions from 1D to 3D models. A series of 1D analyses at different locations within the Tower footprint could be used to
predict tilt related to spatially variable volume change–related settlements. However, for the sake of brevity in this paper, only 3D
models are used to analyze foundation tilt.
The following subsections present the depth-dependent stress
increases considered in the analyses, describe calculations of immediate settlement (from shear distortion), present 1D analyses of
consolidation and secondary compression, and present 3D analyses
of combined volumetric and shear responses.
Stress Increase
The Tower self-weight (Q ¼ 1,055 MN) was derived from column
loads in permit documents. The time rate of self-weight application
(dQ=dt) was developed from documented construction milestones.
Self-weight was reduced in consideration of the excavated soil mass
and the buoyant force from the average groundwater level in Fig. 7.
Fig. 19 shows the resulting time variation of net building load (range
is from Qnet ¼ −75 to 790 MN).
For a specific value of Qnet (corresponding to a particular time),
we considered two methods for analysis of stress increase as a function of depth in the upper OBC soils. Method 1 assumed transfer of
net loads from the piles to the lower marine sand and Colma sand
units through side friction and end bearing (equivalent raft method;
Terzaghi and Peck 1967; Peck et al. 1974; Meyerhof 1976). Those
loads were distributed at a 4:1 slope with depth along the length of
the piles within the lower marine sand and Colma sand units and
then distributed with depth below the equivalent raft according to
the 2:1 method (Hannigan et al. 2006, adapted from Cheney and
Chassie 1993). An average pile tip elevation of −20.6 m (depth of
24.9 m) was used in these analyses. The equivalent raft was taken at
a depth of 23.6 m, which is approximately middepth in the combined marine sand/lower marine sand/Colma sand units. Its lateral
dimensions were 31.7 × 47.5 m. The stress distribution from the
equivalent raft method was found to be broadly consistent with
stress transfer from the piles to the soil as informed by the 3D
model, described subsequently.
Method 2 evaluated stress increase from the 3D model of the
site. Details of the 3D model are presented subsequently (“ThreeDimensional Deformation Analysis” section). This method of analysis included the individual piles (thus avoiding the need for the
equivalent raft approximation) and the horizontally variable stratigraphy of the site. Stress increase from this model depended on horizontal position beneath the building footprint as well as depth.
Within the footprint of the Tower foundation, an area about ⅓–½
of the footprint area located northwest of center generally experienced the largest stress increases (details in Section S2 in the
Supplemental Materials). The total stress increase in this area without effects of adjacent construction was compared to that from
Method 1 in Fig. 16. Because the stress increases from Methods 1
and 2 were similar over the depth range that most strongly contributes to the settlement (i.e., below the especially overconsolidated
crust from 27 to 32 m depth), only a single result (from the equivalent raft method) was used in the 1D analyses presented subsequently in this section. The 3D analyses of settlement used the
spatially variable stress increases derived in those analyses.
Immediate Settlement
We considered three sources of immediate settlement: elastic pile
shortening, immediate settlements in the Colma sand unit, and shear
deformations in the upper OBC unit. Elastic pile shortening above
the equivalent raft was computed from axial strains integrated over
effective pile lengths of 16 m (distance from foundation mat to equivalent raft) and Young’s moduli from ACI (2008) (34,000 MPa).
Immediate settlements in the Colma sand were computed by applying the method of Burland and Burbidge (1985) to the equivalent
raft. These analyses considered the problem geometry, including
the layer thickness below the equivalent raft (3.8 m), and the average
standard penetration test blow count of 29 over this depth range.
Immediate settlement from shear deformation of the upper OBC
was calculated using elastic methods applicable for the case of a
vertical uniform stress applied to a finite area. Using procedures
Fig. 19. Time variation of total and net building load.
© ASCE
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by Mayne and Poulos (1999), we considered the effects of modulus
increase with depth, the location of a rigid layer at limited depth
(taken as 15 m below the top of the upper OBC, corresponding approximately to the depth below which the upper OBC only experiences recompression in the analysis), the rigidity of the foundation,
and the depth of embedment of the equivalent raft.
The 3D analyses considered these same mechanisms, but they
are directly incorporated into the model using the same material
properties.
recompression requires an assessment of the time when OCR ¼ 1.5,
which is denoted tOCR<1.5 .
Using these procedures for analysis of secondary compression
increments, the secondary compression settlement for the entire
stratum was computed as
SSt ¼
H
X
t
X
ΔSSz;t
ð4Þ
z¼0 t¼tOCR<1.5
where ΔSSz;t is from Eq. (3).
This work is made available under the terms of the Creative Commons Attribution 4.0 International license.
One-Dimensional Settlement Analysis
One-dimensional methods of analysis are commonly used to assess
the effects of primary consolidation and secondary compression.
The 1D assumption implies horizontal soil layering and vertical
seepage of pore fluids during consolidation. The analyses produce
time-dependent volumetric strains in the soil column (which are
integrated to settlement) but do not consider the effects of shear
deformations. The effects of adjacent construction are considered
in these analyses through groundwater lowering as reflected in the
“Baseline” model shown in Fig. 7. We applied 1D methods to
evaluate their ability to capture observed foundation responses and
to evaluate the effects of soil parametric variability.
Analysis Approach
The time-dependent consolidation settlement of clay is governed
by the diffusion differential equation, which was recast in one dimension by Terzaghi (1925) as the consolidation differential equation. Analytical solutions to the consolidation equation exist for
simple cases, but those solutions do not consider time variations in
loading, hydrostatic pressures, and coefficient of consolidation that
are important for the present application. Accordingly, we solved
the consolidation equation using a finite-difference recurrence approach. This approach, which adapts recurrence relations developed by Harr (1966) as presented in Section 9.3.2 of Holtz et al.
(2011) to incorporate the aforementioned time-variable quantities,
is described in Section S1 in the Supplemental Materials. The 1D
analyses assumed double drainage at the top and bottom of both the
upper and lower OBC units (relatively granular soils at the interface
of these units were taken as a drainage boundary).
Secondary compression was considered in the analysis both during primary consolidation (t < tp ) and following primary consolidation (t > tp ) in a manner that accounted for the effects of OCR on
the secondary compression strain index (Cαε ) of individual layers.
Section S1.5 in the Supplemental Materials describes how secondary compression was considered simultaneously with primary
consolidation, and Section S1.6 describes how the overconsolidation effect on secondary compression rates was accounted for.
Secondary compression settlement for a layer of thickness Δz
over time interval Δt was computed as
0
ΔSSz;t ¼ Cαεz ðCαεz
=Cαεz ÞΔz log
t þ Δt
t
ð3Þ
where Cαεz = secondary compression strain index at depth z ;
0 =C
Cαεz
αεz = ratio of the reduced secondary compression strain index due to overconsolidation and the secondary compression strain
index at depth z; t = elapsed time since the load was applied that
induced primary consolidation; and t to t þ Δt =time interval for
which the secondary compression increment ΔSSz;t is computed.
0 =C
During initial loading, the ratio Cαεz
αεz is taken as zero for
OCR ≥ 1.5 and is otherwise estimated as described in Section S1.6.
Because secondary compression is neglected for high OCRs during
initial loading, analysis of secondary compression associated with
© ASCE
Soil Properties
Material parameters used in the consolidation and secondary compression analyses are derived from the test data presented in the
“Site Characteristics” section. The specific soil layering and baseline properties used in the analyses are given in Table 2. As shown
in Figs. 3 and 16, baseline profiles for each property were established for the upper and lower OBC units based on the data trends.
Variations of compressibility parameters (i.e., CCε and Crε ) and
rate parameters (cv for normally consolidated and overconsolidated conditions; Cαε ) were considered in formulating 12 alternative parameter sets (referred to as Runs 1–12), as shown in
Table 3. The specific changes in compressibility parameter values
shown in Table 3 are up to approximately one standard deviation
(i.e., COVCcε ≈ 26%, COVCrε ≈ 33%, COVCαε ≈ 32%) based on
the data scatter shown in Fig. 3. Two additional parameter sets were
considered for higher and lower σp0 profiles in the upper OBC, as
shown in Fig. 16 (referred to as Runs 13 and 14). These alternative
Table 2. Consolidation parameter profiles for Baseline 1D analysis
Depth
at center
Layer of layer
no.
(m)
Ccε
Crε
Cαε
1
2
3
4
5
6
7
8
9
10
11
12
13
14
15
16
17
18
19
20
21
22
23
24
25
26
27
28
29
30
31
0.200
0.200
0.200
0.250
0.300
0.300
0.300
0.300
0.300
0.300
0.300
0.300
0.300
0.300
0.300
0.300
0.300
0.300
0.300
0.275
0.250
0.250
0.250
0.250
0.250
0.250
0.250
0.250
0.250
0.250
0.250
0.0200
0.0200
0.0200
0.0250
0.0300
0.0300
0.0300
0.0300
0.0300
0.0300
0.0300
0.0300
0.0300
0.0300
0.0300
0.0300
0.0300
0.0300
0.0300
0.0275
0.0250
0.0250
0.0250
0.0250
0.0250
0.0250
0.0250
0.0250
0.0250
0.0250
0.0250
0.00620
0.00620
0.00620
0.00775
0.00930
0.00930
0.00930
0.00930
0.00930
0.00930
0.00930
0.00930
0.00930
0.00930
0.00930
0.00930
0.00930
0.00930
0.00930
0.00853
—
—
—
—
—
—
—
—
—
—
—
28.2
29.7
31.2
32.8
34.3
35.8
37.3
38.9
40.4
41.9
43.4
45.0
46.5
48.0
49.5
51.1
52.6
54.1
55.6
57.2
58.7
60.2
61.7
63.2
64.8
66.3
67.8
69.3
70.9
72.4
73.9
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Cv;OC
σp0
Cv;NC
2
2
(m =year) (m =year) (kPa)
5.02
4.97
4.93
3.64
3.02
3.02
3.02
3.04
3.06
3.10
3.14
3.19
3.24
3.31
3.38
3.46
3.55
3.64
3.73
4.09
4.52
4.54
4.57
4.60
4.63
4.66
4.69
4.72
4.76
4.79
4.82
37.82
37.82
37.82
26.08
21.73
21.91
22.08
22.43
22.78
23.30
23.82
24.52
25.21
26.08
26.95
27.99
29.04
30.25
31.30
34.14
37.56
37.56
37.56
37.56
37.56
37.56
37.56
37.56
37.56
37.56
37.56
694
694
694
599
599
603
608
618
627
642
656
675
694
718
742
771
800
833
862
862
862
862
862
862
862
862
862
862
862
862
862
J. Geotech. Geoenviron. Eng.
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Table 3. Parameter variations for Runs 1–14 as compared to the Baseline 1D analysis
Analysis run
Ccε
Crε
Cαε
Baseline
1
2
3
4
5
6
7
8
9
10
11
12
13
14
−30%
−20%
−10%
−5%
þ5%
þ10%
þ20%
þ30%
−10%
−5%
þ5%
þ10%
−30%
þ30%
−30%
−20%
−10%
−5%
þ5%
þ10%
þ20%
þ30%
−10%
−5%
þ5%
þ10%
0%
0%
−30%
−20%
−10%
−5%
þ5%
þ10%
þ20%
þ30%
−10%
−5%
þ5%
þ10%
−30%
þ30%
Cv;NC
Cv;OC
σp0 profile
þ30%
þ20%
þ10%
þ5%
−5%
−10%
−20%
−30%
þ30%
þ15%
−10%
−20%
−64% to 66%
þ92% to 99%
þ30%
þ20%
þ10%
þ5%
−5%
−10%
−20%
−30%
þ30%
þ15%
−10%
−20%
−15% to 23%
þ0 to 7.5%
Fig. 16 baseline
Baseline
Baseline
Baseline
Baseline
Baseline
Baseline
Baseline
Baseline
baseline
Baseline
Baseline
Baseline
Low
High
Fig. 3
Fig. 20. Comparison of observed planar average settlement and calculated settlement contributions from 1D model using baseline and variable soil
properties.
parameter sets were formulated in consideration of data scatter,
parameter correlations (e.g., CCε , Crε , and Cαε were varied from
baseline in a similar manner), and degree of realism given data
trends (discussed further subsequently). Variations in the depth
to groundwater over the time horizon of the initial construction
(i.e., through 2007) were considered in additional runs and found
not to significantly influence results; these results are not presented
here for brevity.
Results
Using the approach described previously with the net foundation
loading in Fig. 19 and the groundwater variations in Fig. 7, consolidation, secondary compression, and immediate settlements
were computed for baseline soil properties. These are shown in
Fig. 20, as well as the total predicted settlements and measured settlements (planar average and SM11 and 27 to indicate the range).
Consolidation made up about 70% of the cumulative settlement
as of early 2020, the balance being roughly equally divided among
immediate settlements and secondary compression. Immediate settlements have been essentially constant since 2009. Secondary compression has been a significant portion of the continuing settlement
since early 2019.
The overall settlement amount compares favorably to the planar
average, including the relatively modest rate of settlement since
early 2019. As described in the “Foundation Settlement” section,
© ASCE
measured foundation settlements accelerated at several points in
time coinciding with construction activities (excavation, dewatering)
at adjacent sites. These effects were also evident from the simulations, most notably as downward inflections of consolidation settlements in mid-2012 and the start of 2015. The principal misfit of the
1D simulations from the planar average settlement was an underprediction of settlement rate starting in 2013, at the time of STC Zone
3–4 and SFE dewatering. This underprediction causes simulated settlements to fall below measurements from 2013 to 2018. Potential
causes of underpredicted settlement rates in 2013 include too-slow
consolidation in the 1D model resulting from 3D flow dissipating
excess pore pressures and potentially shear-related movements from
excavations (examined further in the next subsection).
In the “Performance Assessment” section, we inferred the relative contributions of primary consolidation and secondary compression on observed settlement rates. This issue is explored further by
interpreting the excess pore water pressure distribution with depth
from the 1D simulations to evaluate the average degree of consolidation (U 1D ), computed as
U 1D ¼
1
H
Z H
0
Δσz − uz;t
dz
Δσz
ð5Þ
where Δσz = stress change inducing consolidation at depth z
(influenced by Qnet and groundwater changes, as discussed in
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Section S1); and uz;t = excess pore pressure at depth z and time t.
At the end of 2011, which immediately preceded the onset of major STC construction, U 1D was approximately 81% for the upper
OBC and 100% for the lower OBC. This suggests that primary
consolidation may have contributed modestly to the settlement rate
shown in Fig. 15. In April 2019 and May 2020, the 1D simulations
indicated U 1D ¼ 96% and 98%, respectively, for the upper OBC
and U 1D ¼ 100% for the lower OBC at both times, which is consistent with inferences of consolidation having been essentially
completed from both observed settlement rates (Fig. 18) and measured pore pressures (Fig. 8).
The groundwater rebound of 2.1 m that occurred between 2018
and mid-2020 lowered Δσz and has overconsolidated portions of
the upper OBC, which previously experienced virgin compression.
The effective OCR of this sublayer was 1.02 in May 2020 in the 1D
analysis, which slows secondary compression by reducing Cαε . This
effect is included in the computed settlements shown in Fig. 20.
Fig. 20 also shows variations of total settlements across the 15
runs reflecting soil parametric variability. The variability among the
settlement results was modest, which is somewhat by design—the
variations shifted compressibility parameters and cv in opposite directions relative to baseline (e.g., Run 1 had lower compressibility,
higher cv ; Run 2 had higher compressibility, lower cv ), which was
done to reflect correlations between these parameters. The results
show that reasonable variations in soil compressibility parameters
and cv did not affect the ability of 1D methods to capture the main
settlement features. On the other hand, uniform modifications up or
down of preconsolidation pressures across the OBC profile (not
included in the 15 variability runs) shifted down and up, respectively, computed settlements. Such results are not shown in Fig. 20.
As described previously, the σp0 profile used in the analysis
(Fig. 16) was selected to represent trends in laboratory data while
also producing computed settlements that are consistent with measurements. Variations in the σp0 profile were considered for the
upper OBC as shown in Fig. 16. If these σp0 profile changes were
applied while maintaining all other properties at baseline levels,
settlements would shift up and accelerate for lower σp0 and shift
down and decelerate for higher σp0 , producing mismatches with observation. Instead, we modestly adjusted compressibility and rate
parameters along with σp0 , as shown in the last two rows of Table 3,
to see if reasonable agreement with data could be obtained. For the
lower σp0 case (Run 13), even with reduced compressibility parameters (within the range of data in Fig. 3), calculated settlements exceeded observations and had too-high settlement rates in mid-2020
(indicating continuing consolidation). This indicates that the lower
σp0 profile cannot reasonably replicate field performance when used
in 1D analyses. For the higher σp0 case (Run 14), the calculated
settlement approximately matched the lower bound observed settlement for the initial loading and postconstruction time period.
However, the σp0 profile was too high for significant virgin compression to occur, which produced mismatched settlement rates, especially from dewatering (since 2012).
Three-Dimensional Deformation Analysis
Three-dimensional methods of analysis have the ability to account
for several factors that are neglected in 1D analyses, including spatial variations in soil layer thicknesses and properties, spatial variations in water table depths, more realistic (nonvertical) seepage
paths, and shear deformations. The effects of adjacent construction
were considered in these analyses through groundwater lowering
(as in the 1D analyses) as well as through shear deformations from
stress changes (e.g., excavations). The objective of the 3D analyses
© ASCE
described here was to qualitatively assess the significance of these
effects on the computed foundation performance. For brevity, the
results reported here considered only baseline soil parameters used
in 1D analyses; no “tuning” of parameters was applied to improve
the fit of computed foundation deformations to measurements. Moreover, no randomization of stress history or compressibility parameters in the horizontal plane was considered.
Analysis Approach
FLAC3D version 7 (Itasca Consulting Group, Inc. 2019) was used
to model the soil–structure interaction, hydromechanical coupling,
and time-dependent (viscous) elastoplastic behavior of the soil in
the overall system. The Supplemental Materials (Section S2.1) describe details of the analyses, including the sequencing of construction phases in the model.
As in the 1D analyses, variations with time of net foundation
loading were considered as shown in Fig. 19. Dewatering of adjacent sites was included in the model, and groundwater levels in the
model domain were allowed to fluctuate in response to the dewatering (see Section S2.3).
Material Models
Three constitutive models were used for different materials: elastic,
plastic hardening (PH; Schanz et al. 1999), and soft soil creep (SSC)
models (Stolle et al. 1999; Vermeer and Neher 1999). The fundamental soil properties used to set model parameters were taken from
the baseline profiles in Fig. 3 and matched those used in the 1D
analyses. Section S2.2 presents details on implementation of the
material models, a calibration process undertaken for the SSC model
in OBC, modeling of the Tower’s pile foundations, and the specification of soil hydraulic conductivity parameters.
Results
The FLAC3D simulations captured consolidation, secondary compression, shear deformations, and immediate settlements. Fig. 21
plots the resulting settlement versus time along with the planar
average settlement and the computed settlement from 1D analysis.
The 3D analyses produced similar amounts of cumulative settlement and matched the observations well. The similarity of the
1D and 3D cumulative settlements suggests that volume change
was the dominant mechanism causing settlements; this is inferred
because volume change is considered in both sets of analyses
whereas shear deformations are only considered in the 3D case.
Relative to the 1D analyses, the 3D analyses better represented
the settlement inflection point in early 2013. The modeled settlements rebounded in 2019, which was not seen in the observed settlements. This rebound was due to groundwater in the model
recharging more quickly than the field measurements.
To estimate the degree of consolidation from 3D analysis for a
given time (U 3D ), we computed stress change and excess pore pressures for the finite-difference zones within the envelope of the
building foundation across the thickness of the OBC layers. The
average stress change and excess pore pressure for a given depth
were computed, the profiles of which were then used to estimate
U 3D , as in Eq. (5). At the end of 2011, U 3D ¼ 90% in the upper
OBC and 100% for the lower OBC. The 90% consolidation in the
upper OBC was higher than U 1D ¼ 81% for this same time, due to
allowance for horizontal seepage. The higher degree of pore pressure dissipation in the 3D analyses caused the inflection point in
2013 (when additional pore pressures were generated from adjacent
construction) to be more pronounced, which in turn produced better
agreement with the data for the time period 2013–2018. At the start
of 2019, the 3D simulations indicated U 3D was about 99% in the
upper OBC, which again supports the interpretation of consolidation being essentially complete.
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Fig. 21. Tower mat settlement comparison among measurements, 1D analyses, and 3D analyses.
Fig. 22. Comparison of observed horizontal roof deflections with calculated horizontal roof deflections from 3D model using baseline soil properties:
(a) north–south deflections; and (b) east–west deflections.
Fig. 22 compares measured horizontal roof deflections from
foundation tilt to computed results from the 3D model (these deflections were zero for the 1D model). Unlike measured deflections, computed deflections started during Tower construction, and
the values at the onset of measurements (early 2009) were 10 cm to
the north and 15 cm to the west, which can be compared to dataderived estimates (Figs. 12 and 13) of the deflections at that time of
10–12 cm (north) and 1–8 cm (east). Since 2009, the 3D model
overestimated both cumulative deflection and change in deflection
to the north by about 10 cm. The 3D model overestimated the
cumulative deflection to the west by about 18 cm, but the change
in deflection since 2009 (42 cm) compared favorably with measurements (44 to 50 cm).
© ASCE
In the north–south direction, the effects of adjacent construction
were evident in the computed deflections in a similar manner to that
shown by the data, including:
• Reversal of deflection toward the south in 2012;
• Reversal of deflection toward the north in 2013; and
• Relatively small change in deflection from 2014 to 2020.
In the east–west direction, the effects of adjacent construction
were not readily apparent in the computed deflections. Whereas
the measured data showed clear inflection points in 2013, 2015,
and 2017, the slope of the computed deflections was relatively
constant. Computed deflections were influenced by too-large moments applied to the foundation from P-delta effects associated
with overpredicted westerly deflections in 2009. Our interpretation
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is that these moments may have produced relatively large tilt deflections from consolidation and shear of foundation soils, reducing sensitivities to groundwater variations, as occurred in the real
structure.
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Summary and Conclusions
This paper documents and analyzes the Millennium Tower case history of foundation movement over the time period 2006–2020. The
value of this case history stems from the detailed information on
foundation movements that can be used to interpret controlling mechanisms and to validate predictive models of consolidation or secondary compression and foundation movements from adjacent
excavations. The objectives of the paper are to document the case
history, interpret and analyze the data to elucidate deformation mechanisms in the foundation soils that led to the foundation movements,
and evaluate whether the observed movements are predictable.
We describe site stratigraphy, key attributes of which are sand
units extending approximately from 10 to 24 m depth and an underlying overconsolidated marine clay (upper Old Bay Clay). Compressibility and strength parameters of OBC were established from
laboratory testing. We describe the Millennium Tower structure, its
adjoining podium, and a series of neighboring structures mainly
constructed from 2012 through mid-2018. We found that a critical
aspect of the construction of these adjacent structures was dewatering and excavation for construction of substructures. Finally, we
describe measured data from instrumentation, which included pore
pressure measurements from piezometers, foundation and Tower
movements (both settlement and tilt), and inclinometers measuring
horizontal ground deflections.
Groundwater levels at the Tower site were lowered appreciably
during podium construction but recovered rapidly starting in early
2008 once on-site dewatering was discontinued. Piezometric data
show that adjacent construction lowered ground water at the Tower
site in multiple episodes from mid-2012 to mid-2018, one of which
(STC) was maintained for 6 years. The Tower settled approximately
40 cm in total, approximately half of which has occurred since the
completion of Tower construction (late 2009). At roof level, the
Tower has tilted 40 cm in the project west direction (since late
2009) and 17 to 18 cm in the project north direction (which has
only modestly increased since late 2009). The Tower foundation
movements in both settlement and tilt have exhibited inflections
at various times since Tower construction coinciding with dewatering and excavation.
Engineering analysis of the case history with 1D and 3D methods demonstrates that the observed settlements were predictable in
terms of cumulative amounts and their time variations. The settlements were dominated by volumetric deformations in the OBC unit
from primary consolidation and secondary compression. Portions
of this unit over the approximate depth range 33–45 m experienced
virgin compression, which was the principal contributor. The role
of primary consolidation has changed over time, dominating during
construction at the Tower site (2005–2009) and from approximately
2012–2018. On the other hand, secondary compression was a significant contributor in the period immediately preceding adjacent
construction (2010–2011) and has dominated from 2019 onward.
Computed horizontal deflections from 3D analyses were able to capture the general directions and inflection points from observations in
the project north–south direction, but did not capture the cumulative
deflection amounts nor the inflection points in the project east–west
direction.
This study is not the first to document long-term foundation settlement and tilt of a tall structure. For example, prior case histories
© ASCE
of consolidation settlement of tower structures were presented by
Briaud et al. (2007, 2009, 2015), which demonstrated the combined
effects of tower weight and regional groundwater drawdown on
settlement patterns. Moreover, case studies of ground movements
adjacent to supported excavations were presented by Peck (1969),
Long (2001), Moormann (2004), Korff (2013), and Korff et al.
(2016), which demonstrated that the largest settlements occurred
near excavations and the amount of settlement generally decreased
with distance, thus producing tilt toward the excavation. These deformation patterns are consistent with those documented here for the
Millennium Tower.
Data Availability Statement
Data on consolidation and secondary compression properties of the
upper OBC soils was presented in Wagner et al. (2021) and its supplemental documents. The other data provided in this case history
paper were assembled from professional reports that are cited as
well as the following documents:
• Arup North America, Ltd. (2009). Transbay Transit Center:
Results of Settlement Survey at 301 Mission Property, data
transmission.
• Arup North America, Ltd. (2010). Correspondence from Nick
O’Riordan Re: Survey of PG&E Vault Adjacent to 301 Mission
Tower, September 24.
• Arup North America, Ltd. (2015–2017). Salesforce Tower
Monitoring Reports—No. 1 through 37, January 29, 2015
through August 23, 2017.
• Langan Engineering (2017–2020). Building Monitoring Report(s)
for 301 Mission Street, data transmissions.
• Martin M. Ron Associates (MM Ron) (2006–2009). Summary
of Surveyed Settlements: 301 Mission, data transmission.
The data from these sources have been assembled into supplemental tables to accompany the figures used in the paper. The publicly available documents can be obtained by contacting the San
Francisco Department of Inspections (dbi.ppcrequest@sfgov.org)
with the block or lot numbers for the project area (accessible here:
https://sfplanninggis.org/pim/).
Acknowledgments
The authors thank Mission Street Development, LLC, which was the
developer for the Millennium Tower project, for funding the work
presented here. The first author was a presenting expert in legal proceedings related to causation. Our objectives in presenting this paper
are as stated in the “Introduction” and are unrelated to the interests of
any party in the legal proceedings. Many individuals and groups
contributed both directly and indirectly to this effort in a variety of
capacities, including Ron Hamburger, Lachezar Handzhiyski, and
others from Simpson Gumpertz and Heger, Inc.; Sean Culkin from
Integral Consulting, Inc.; Sydney Maguire, Barry Zheng, and Tom
Clifford from Slate Geotechnical Consultants, Inc.; various consultants at Langan Engineering and Environmental Services; and various consultants at ENGEO, Inc. We thank the two sets of three
anonymous reviewers for different versions of this paper, whose
constructive feedback has improved the paper.
Supplemental Materials
The supplemental materials include a detailed plan view of the site
showing locations of subsurface exploration (Plate S1); a spreadsheet file containing data from selected figures, geodetic coordinates
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J. Geotech. Geoenviron. Eng., 2023, 149(6): 05023002
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of subsurface explorations (borings, CPTs, etc.), and selected boring
logs; and a document describing foundation performance simulations and containing Figs. S1–S8, Tables S1–S3, and Eqs. (S1)–
(S22). These materials are available online in the ASCE Library
(www.ascelibrary.org).
This work is made available under the terms of the Creative Commons Attribution 4.0 International license.
References
ACI (American Concrete Institute). 2014. Building code requirement for
structural concrete and commentary. ACI 318-12. Farmington Hills,
MI: ACI.
Arup North America, Ltd. 2009. Report on the preliminary design of the
buttress for limiting the excavation-induced deformations, 2009–2019.
San Francisco: Transbay Transit Center Project.
Arup North America, Ltd. 2010. Transbay Transit Center program: Transbay Transit Center. Contract No. 08-04-CMGC-000. Los Angeles:
Arup North America.
Arup North America, Ltd. 2013. Transbay Tower final geotechnical
data report. Job Number 229478-00. Los Angeles: Arup North
America.
Arup North America, Ltd. 2018. Memorandum, Salesforce Transit Center,
supplemental instrumentation installation. Los Angeles: Arup North
America.
Bjerrum, L. 1967. “Engineering geology of Norwegian normallyconsolidated marine clays as related to settlement of buildings.” Géotechnique 17 (2): 83–118. https://doi.org/10.1680/geot.1967.17.2.83.
Brandenberg, S. J. 2017. “iConsol.js: JavaScript implicit finite-difference
code for nonlinear consolidation and secondary compression.” Int. J.
Geomech. 17 (6): 04016149. https://doi.org/10.1061/(ASCE)GM.1943
-5622.0000843.
Briaud, J. L., Y. Koohi, J. Nicks, and I. Jung. 2015. “San Jacinto monument: New soil data and analysis including subsidence.” J. Geotech.
Geoenviron. Eng. 141 (6): 04015023. https://doi.org/10.1061/(ASCE)
GT.1943-5606.0001299.
Briaud, J. L., J. Nicks, K. Y. Rhee, and G. Stieben. 2007. “San Jacinto monument case history.” J. Geotech. Geoenviron. Eng. 133 (11): 1337–1351.
https://doi.org/10.1061/(ASCE)1090-0241(2007)133:11(1337).
Briaud, J. L., B. Smith, K. Y. Rhee, H. Lacy, and J. Nicks. 2009. “The
Washington monument case history.” Int. J. Geoeng. Case Hist. 1 (3):
170–188.
Burland, J. B., and M. C. Burbidge. 1985. “Settlement of foundations on
sand and gravel.” Proc. Inst. Civ. Eng. 78 (6): 1325–1381. https://doi
.org/10.1680/iicep.1985.1058.
CBS (Columbia Broadcasting System). 2017. “San Francisco’s leaning
tower of lawsuits.” Accessed January 30, 2017. https://www.cbsnews
.com/news/millennium-tower-san-francisco-leaning-tower-of-lawsuits-60
-minutes/.
Cheney, R. S., and R. G. Chassie. 1993. Soil and foundation workshop
manual. 1st ed. Washington, DC: Federal Highway Administration.
Dames & Moore. 1966. Boring logs for 50 Beale Street, San Francisco,
California, prepared for Bechtel Corporation San Francisco Office
Building by Skidmore, Owings & Merrill (SOM). Los Angeles:
Dames & Moore.
Dames & Moore. 1981. Geotechnical investigation, proposed Five Fremont
Center project, San Francisco, California, prepared for Metropolitan
Bechtel Shorenstein. Los Angeles: Dames & Moore.
DeSimone Consulting Engineers, LLC. 2005. Foundation permit submittal
volume I—Project overview and structural analysis, Volume II—
Foundation design, 301 Mission Street, San Francisco, CA. New York:
DeSimone Consulting Engineers.
Hamburger, R. O., S. E. Harris, and L. V. Handzhiyski. 2021. “Stabilizing
San Francisco’s leaning tower.” In Structure, 8–10. Chicago: National
Council of Structural Engineers Associations.
Hannigan, P. J., G. G. Goble, G. E. Likins, and F. Rausche. 2006. Design
and construction of driven pile foundations. Rep. No. FHWA-NHI-05042. Washington, DC: Federal Highway Administration.
Harr, M. E. 1966. Foundations of theoretical soil mechanics, 381. New York:
McGraw-Hill.
© ASCE
Holtz, R. D., W. D. Kovacs, and T. C. Sheahan. 2011. An introduction to
geotechnical engineering. 2nd ed. London: Pearson.
Itasca Consulting Group, Inc. 2019. FLAC3D 7.0 user’s manual.
Minneapolis: Itasca.
Kluzniak, B., D. Koutsoftas, and N. O’Riordan. 2010. “Prototype test of soilcement shoring walls for the Transbay Transit Center, San Francisco.”
In Proc., Earth Retention Conf. 3, 294–302. Reston, VA: ASCE.
Korff, M. 2013. “Response of piled buildings to the construction of
deep excavations.” Ph.D. thesis, Dept. of Engineering, Univ. of
Cambridge.
Korff, M., R. J. Mair, and F. A. Van Tol. 2016. “Pile–soil interaction and
settlement effects induced by deep excavations.” J. Geotech. Geoenviron. Eng. 142 (8): 04016034. https://doi.org/10.1061/(ASCE)GT.1943
-5606.0001434.
Kutter, B. L., and N. Sathialingam. 1992. “Elastic-viscoplastic modelling of
the rate-dependent behaviour of clays.” Géotechnique 42 (3): 427–441.
https://doi.org/10.1680/geot.1992.42.3.427.
Ladd, C. C., and R. Foote. 1974. “A new design procedure for stability of
soft clays.” J. Geotech. Eng. Div. 100 (7): 763–786. https://doi.org/10
.1061/AJGEB6.0000066.
Long, M. 2001. “Database for retaining wall and ground movements due to
deep excavations.” J. Geotech. Eng. Div. 127 (3): 203–224. https://doi
.org/10.1061/(ASCE)1090-0241(2001)127:3(203).
Mayne, P. W., and H. G. Poulos. 1999. “Approximate displacement influence factors for elastic shallow foundations.” J. Geotech. Geoenviron.
Eng. 125 (6): 453–460. https://doi.org/10.1061/(ASCE)1090-0241(1999)
125:6(453).
Mesri, G., and Y. K. Choi. 1985. “The uniqueness of the end-of-primary
(EOP) void ratio-effective stress relationship.” In Proc., 11th Int. Conf. on
Soil Mechanics and Foundation Engineering, 587–590. Rotterdam,
Netherlands: A.A. Balkema.
Meyerhof, G. G. 1976. “Bearing capacity and settlement of pile foundations.” J. Geotech. Eng. Div. 102 (3): 197–228. https://doi.org/10.1061
/AJGEB6.0000243.
Moormann, C. 2004. “Analysis of wall and ground movements due to deep
excavations in soft soil based on a new worldwide database.” Soils
Found. 44 (1): 87–98. https://doi.org/10.3208/sandf.44.87.
Peck, R. 1969. “Deep excavation and tunneling in soft ground.” In Vol. 7
of 7th Int. Conf. on Soil Mechanics and Foundation Engineering,
225–290. Mexico City: Sociedad Mexicana de Mecanica.
Peck, R. B., W. E. Hanson, and T. H. Thornburn. 1974. Foundation engineering. 2nd ed. New York: Wiley.
SAGE Engineers, Inc. 2018. Data report for geotechnical investigation and
piezometer installation, 301 Mission Street, City and County of San
Francisco, California. Westfield, MA: SAGE Engineers.
Schanz, T., P. A. Vermeer, and P. G. Bonnier. 1999. “The hardening soil
model: Formulation and verification.” In Proc., Int. PLAXIS Symposium:
Beyond 2000 in Computational Geotechnics, edited by R. B. J.
Breinkgreve, 281–296. Oxfordshire, UK: Taylor & Francis.
Slate Geotechnical Consultants, Inc. 2019. Geotechnical investigation &
instrumentation installation—Plenum area, 301 Mission Street, San
Francisco, California. South Lake Tahoe, CA: Slate Geotechnical
Consultants.
Stolle, D. F. E., P. A. Vermeer, and P. G. Bonnier. 1999. “Time integration
of a constitutive law for soft clays.” Commun. Numer. Methods Eng.
15 (8): 603–609. https://doi.org/10.1002/(SICI)1099-0887(199908)
15:8<603::AID-CNM274>3.0.CO;2-F.
Taylor, D. W., and W. Merchant. 1940. “A theory of clay consolidation accounting for secondary compression.” J. Math. Phys. 19 (1–4): 167–185.
https://doi.org/10.1002/sapm1940191167.
Terzaghi, K. 1925. “Erdbaumechanik auf Bodenphysikalischer Grundlage,
Franz Deuticke, Leipzip und Wein.” In From theory to practice in soil
mechanics: Selections from the writings of Karl Terzaghi with bibliography and contributions on his life and achievements, 10–13. Translated by A. Casagrande. New York: Wiley.
Terzaghi, K., and R. B. Peck. 1967. Soil mechanics in engineering practice.
2nd ed. New York: Wiley.
Terzaghi, K., R. B. Peck, and G. Mesri. 1996. Soil mechanics in engineering practice. 3rd ed. New York: Wiley.
05023002-21
J. Geotech. Geoenviron. Eng., 2023, 149(6): 05023002
J. Geotech. Geoenviron. Eng.
Treadwell & Rollo, Inc. 2012. Geotechnical investigation, 350 Mission
Street, San Francisco, California. Project No. 730466502. San Francisco:
Treadwell & Rollo.
Vermeer, P. A., and H. P. Neher. 1999. “A soft soil model that accounts for
creep.” In Proc., Int. PLAXIS Symp.: Beyond 2000 in Computational
Geotechnics, edited by R. B. J. Breinkgreve, 249–261. Rotterdam,
Netherlands: A.A. Balkema.
Wagner, N., M. Largent, J. P. Stewart, C. Beyzaei, D. Murphy, J. Butkovich,
and J. A. Egan. 2021. “Stress history-dependent secondary compression
of San Francisco Bay region Old Bay clays.” J. Geotech. Geoenviron.
Eng. 147 (7): 04021045. https://doi.org/10.1061/(ASCE)GT.1943-5606
.0002525.
This work is made available under the terms of the Creative Commons Attribution 4.0 International license.
TJPA (Transbay Joint Powers Authority). 2019. “Transbay Joint Powers
Authority board meeting(s).” Accessed January 30, 2017. https://tjpa
.org/news-information/calendar/board-meeting-dates.
Treadwell & Rollo, Inc. 1997. Geotechnical investigation, 350 Mission
Street Building Seismic Strengthening, San Francisco, California.
Project No. 2152.01. San Francisco: Treadwell & Rollo.
Treadwell & Rollo, Inc. 2005. Revised geotechnical investigation 301
Mission Street, San Francisco, California. Project No. 3157.02.
San Francisco: Treadwell & Rollo.
Treadwell & Rollo, Inc. 2006. Correspondence from Christopher Ridley
Re: Summary of pile driving, 301 Mission Street, San Francisco,
California. San Francisco: Treadwell & Rollo.
© ASCE
05023002-22
J. Geotech. Geoenviron. Eng., 2023, 149(6): 05023002
J. Geotech. Geoenviron. Eng.