FOUNDATION CONSTRUCTION CHALLENGES AT 100 11th AVENUE IN MANHATTAN – SECANT PILE WALL

Andrew Cushing and Stephen Young, Arup, New York, NY, USA Alfredas Daugiala and Fabio Liscidini, Underpinning and Foundation Skanska, New York, NY, USA

INTRODUCTION Arup was retained by Underpinning and Foundation Skanska to design a secant pile wall as excavation support for a proposed two-level basement of a new 23-story residential building located at 100 11th Avenue and West 19th Street in Manhattan. The site is bounded to the north by 5 to 8-story structure supported on concrete-filled, tapered corrugated steel casing (ie, Raymond “Step-Taper” piles) and to the east by a two-story building supported on shallow footings of unknown dimension. The western boundary of the site (11th Avenue / West Side Highway) is skewed, resulting in a trapezoidal-shaped site as shown in Figure 1. The initial temporary works design called for drilled pile underpinning of the two-story structure and the use of a relatively flexible temporary sheet pile wall as earth support. Underpinning and Foundation Skanska proposed the use of a stiffer excavation support system consisting of an internally-braced perimeter secant wall to eliminate the underpinning requirement for the two-story structure. The secant pile wall will also act as the permanent retaining wall for the new building and will support the vertical load of the basement structure and the perimeter column

load of the superstructure. This alternate proposal was subsequently accepted. Since the site is underlain by a thick deposit of very soft organic silt and clay, the potential for excavation-induced ground movements and their impact on the adjacent two-story structure on shallow footings were of the utmost concern.

This paper addresses the design and construction considerations associated with this secant wall, as well as installation details.

This paper presents the design and construction of the foundation for a major high-rise real estate development on the west side of Manhattan. Arup was retained by Underpinning and Foundation Skanska to design a secant pile wall as excavation support for a proposed two-level basement of a new 23-story residential building located at 100 11th Avenue and West 19th Street. The secant pile wall will also act as the permanent basement retaining wall for the new building. The site is bounded on two sides by existing structures and is underlain by a thick deposit of very soft organic silt and clay. Therefore, the potential for ground movements and their impact on an adjacent structure bearing on shallow footings were of the utmost concern. This paper addresses the design and construction considerations associated with this secant wall, as well as installation details. Measured wall movements are also used to assess the validity of the pre-construction soil strength and stiffness design parameters.

Figure 1. Site Plan and Dimensions

Measured wall movements are also used to assess the validity of the pre-construction soil strength and stiffness design parameters. GEOLOGICAL CONDITIONS The ground surface prior to excavation activities was at approximate +1.4mEL (+4.5ftEL) relative to Borough President of Manhattan Datum. A total of 21 test borings had been advanced at the site in 2000 and 2005, as shown in Figure 2. On the basis of these test borings, a typical design geologic section of the site was developed as shown in Figure 3, consisting of approximately 4.6m (15ft) of fill over a 13.7m (45ft) thick layer of very soft organic silt and clay, which is itself underlain by about 6m (20ft) of sand. On the basis of the test borings, dark grey slightly-weathered mica schist bedrock is located between 22 and 27m (72 and 87ft) below ground surface. Groundwater is at approximate +0mEL and is governed by the Hudson River one block to the west of the site. The basement excavation was originally intended to consist of two levels, with an excavation depth of approximately 6.7m (22ft). Subsequent design changes resulted in a more shallow one level basement excavated to 3.4m (11ft) below ground surface to -2mEL (-6.5ftEL). Regardless of the number of basement levels, the material which was to be retained was primarily the surficial fill, which varied from as much as 1.5 to 9m (5 to 30 ft) in thickness across the site. For both basement scenarios, the layer of utmost concern was the underlying 13.7m (45ft) thick very soft organic silt and clay layer and its impact on excavation-induced ground movements and wall bending moments. This layer consisted of Standard Penetration Test (SPT) blow counts in the range of 0 to 6 (primarily 0 to 2). Consultation of the Sanitary Topographical Map of Manhattan (Viele, 1865) indicated that the subject site lies within the area reclaimed from the adjacent Hudson River. To better characterize this very soft layer, cone penetration (CPT) tests were performed at the site. A number of these tests encountered refusal in the fill and were terminated early. Cone tip resistance (qc) data from the two profiles advanced through the entire organic silt and clay layer are provided in Figure 4. These data demonstrate the variability in fill thickness relative to the adopted average value used for

design. They also indicate that the underlying organic silt and clay layer ranged from lightly-overconsolidated (OCR >1) to possibly underconsolidated (OCR < 1). The adopted design undrained shear strength (su) profile in the organic silt and clay had assumed normally-consolidated (OCR = 1) conditions with su = 0.22 σ’vo, giving su in the range of 12 to 24 kPa (250 to 500 psf) between the top and bottom of this very soft layer. This is in general agreement with an su profile obtained by dividing qc by an Nk of 17.

Figure 3. Typical Design Geologic Section

Figure 2. Site Plan With Boring Locations

(Langan, 2005)

A summary of the soil strength and stiffness parameters which were adopted for the design of the secant pile wall is provided in Table 1.

It should be noted that the soil modulus (Es) for the very soft organic silt and clay layer was estimated as 250 times the undrained shear strength (250 su). The values of Es and friction angle (φ’) for the fill and underlying sand layers

Table 1. Soil Design Parameters

Stratum γbulk

(kN/m3)

φ’ su

(kPa)

Es

(MPa)

Fill 19 30° - 10

Silt/Clay 16 20° 12 to 24 3 to 6

Sand 20 35° - 25

Rock 27 44° - 140

were estimated from correlations with the SPT and CPT tests.

For the very soft organic silt and clay, two analyses were performed: a drained analysis based upon effective stresses with φ’ = 20° (c’ = 0), and an undrained analysis with the soil strength characterized by su (with φ = 0°). The analyses indicated that the undrained condition was more critical in evaluating the performance of the wall.

SECANT PILE WALL ANALYSIS & DESIGN Software

The software program OasysFREW (Flexible Retaining Wall Analysis) was used to perform the analysis of the secant pile wall. This program permits the user to consider variations in soil strength and stiffness, wall bending stiffness (EI), strut stiffness and preload, multiple excavation stages, and vertical surcharges. At each construction stage, the program calculates and displays strut loads and profiles of wall bending moment, shear force, and lateral displacement. In addition, the induced lateral earth pressure profiles on each side of the wall (relative to theoretical active and passive limits) are also reported so that the user is aware of the depth and degree to which the ultimate passive resistance of the embedded portion of the wall is mobilized. The program is limited to a two-dimensional, plane strain analysis.

Secant Pile Wall Details

The secant pile wall was to be installed around

Figure 4. Cone Tip Resistance

the entire perimeter of the site, and each secant pile was 1m (3.3ft) in diameter. A guide wall (as shown in Figure 5) was to be installed as a template for both primary (unreinforced) and secondary (reinforced) secant pile installation. Each intermediate secondary pile was reinforced with a W24x306 core beam (or equivalent built-up section), with a center-to-center spacing of 1.67m (5.5ft). Pairs of shear studs were tack-welded to the exterior of both flanges at 300mm

(12inch) intervals. The resulting composite bending stiffness (EI) of the wall (neglecting the unreinforced primary secant piles) was 915,000 kN-m2/m (675,000 kip-ft2/ft). A photograph of two W24x306 core beams prior to installation is shown in Figure 6.

Column Loading

As stated in the introduction to this paper, it was decided that the perimeter secant wall would be designed to support selected eccentric superstructure column loads along the north and east walls. The single maximum column load was 9,080 kN (2,040 kips).

Each core beam which contributed to vertical support of the superstructure was socketed approximately 1m (3.5ft) into the underlying bedrock. All other core beams which did not contribute to vertical structural support were socketed 150mm (6inches) into the bedrock.

Existing Structure and Construction Surcharges

The selfweight of the two story structure on shallow footings along the eastern site boundary was modeled by two superimposed surcharges. The calculated wall load of 88kN/m (6 kips/ft) was distributed as a pressure of 72kPa (1.5 kips/ft2) to an assumed 1.2m (4ft) wide continuous footing, with a 500mm (1.6ft) horizontal offset between the back of the secant wall and the outer edge of retained footing. Behind this continuous wall footing, a global uniform pressure of 14 kPa (300 psf) was applied to simulate the slab loading.

Since the 5 to 8-story building to the north was supported on Raymond Step-Taper piles, no external surcharge was applied for this building.

A temporary construction surcharge of 28 kPa (600 psf) of 6m (20 ft) width was applied directly behind the western and southern walls at ground level.

Bracing Plan and Strut Forces

The as-built temporary lateral support for the secant wall consisted of diagonal bracing (W sections) in each of the four corners of the site, in addition to two cross-lot pipe struts (760mm diameter, 25mm wall thickness) supporting the

Figure 6. W24x306 Core Beams

Figure 5. Secant Pile Guide Wall

north and south walls. An approximate bracing plan showing the as-built condition is provided in Figure 7.

The strut stiffness is uniformly-distributed along the length of the wall. Typical values of strut stiffness used in the analysis are summarized as follows:

• Diagonal: 160 MN/m/m (3,300 kip/ft/ft)

• Cross-lot: 90 MN/m/m (1,900 kip/ft/ft)

The calculated stiffness of the diagonal struts is approximately 75% larger than the cross-lot struts. In addition, a diagonal strut positioned within the corner of a square or rectangular excavation is inherently stiffer than a strut supporting a long continuous section of excavation, a phenomenon attributed to field divergence from the typical (conservative) two-dimensional plane strain design assumption.

Analysis Output – Undrained (Total Stress) Analysis

Sample undrained (total stress) analysis output is provided in Figure 8 for the eastern wall (retaining the two story building on shallow footings) for the temporary condition in which the excavation has just reached a maximum depth of approximately 3.4m (11ft). Figure 8 shows profiles of bending moment, shear force, and lateral wall displacement with depth. In addition, a maximum lateral wall displacement of 75mm (3 inches), as well as a

lateral bracing force of 311 kN/m (21 kips/ft), had been predicted for this wall. The predicted displacement was much higher than what would be expected for such a shallow excavation, but it reflects the building surcharge and the presence of a relatively thick layer of very soft organic silt and clay. SITE WORK AND WALL CONSTRUCTION

The site had been used as an underground gas storage facility by the Consolidated Gas Company until 1921, after which time it had served as an auto repair shop, motor freight station, office facility, and most recently as a parking lot. A 27m+ (90ft+) diameter buried brick tank with a 600mm (2ft) thick wall was situated near the center of the site as shown in Figure 2 to a depth of approximately 7.5m (25ft) below grade. The vertical support for this underground tank most likely consists of a raft foundation of concrete or brick and timber planks.

Initial excavation work consisted of digging a trench approximately 900mm (3ft) deep and 1.6m (5.3ft) wide at the base for the installation of the guide wall (Figure 5) for secant pile installation. The parking lot asphalt and subbase was removed and replaced with compacted crushed stone to serve as a construction pad for both secant pile and interior foundation pile installation.

Two secant pile installation rigs were used onsite – a Bauer BG28 and a Casagrande C600.

Figure 8. East Wall - Forces, Moments, and Displacements

Figure 7. Approximate As-Built Bracing Plan

Temporary 1m (3.3ft) diameter steel casing was employed to support the hole through the zones of saturated fill and underlying sand, in addition to the very soft organic silt and clay layer.

During excavation process, the buried brick gas holder tank had to be deconstructed and removed. Remnants of this brick tank, along with a portion of the northern secant wall and interior micropiles, can be seen in Figure 9. Additional secant pile and bracing photographs are provided in Figures 10 through 12.

Figure 10. Bracing of North Wall

Figure 12. Secant Pile Core Beams and Bracing

Figure 11. West Wall Core Beams

(Looking South)

Figure 9. Remnants of Buried Brick Gas Tank

Along North Wall

MEASURED WALL MOVEMENT AND SUBSEQUENT BACK-ANALYSIS

A total of three inclinometers were installed within the secant wall at approximate midpoints of the western, northern, and eastern sides. Unfortunately, the inclinometer installed on the critical eastern wall malfunctioned and did not produce useable results. However, movement data from the other two inclinometers were available for back analysis.

According to the LMW Engineering Group (2007), the inclinometers were zeroed in August 2007 and measured on October 26, 2007. On this latter date, the bracing had been installed across the entire site at approximate -0.3mEL (-1ftEL). To the best knowledge of the authors (and on the basis of site photographs taken on October 20, 2007), no further excavation had taken place along the west wall on October 26, while excavation near the middle of the north wall was nearing the final proposed base of excavation at -2mEL (-6.5ftEL). Furthermore, it should be noted that the construction surcharge used in the design stage for the west wall was not present.

On the basis of this knowledge, the results of a back–analysis of the inclinometer data along the west wall are presented in Figure 13. The measured inclinometer data for wall movement into the excavation is compared to the back analysis using the original soil strength and stiffness parameters summarized in Table 1 (with the construction surcharge removed). The shapes of the two curves are similar in nature, and the maximum measured wall movement of 20mm (0.8 inch) agrees rather closely with the maximum back-analyzed value of 28mm (1.1 inch). The authors do not believe the stiffer-than-expected wall response was the result of conservative soil strength and stiffness parameters. Rather, field divergence from the plane-strain analysis assumption in the analysis program is a more likely explanation for this observed difference.

The results of a similar back–analysis of the inclinometer data along the north wall is presented in Figure 14. It is noted that the rockline is somewhat higher than the average level used at the design stage at this inclinometer location, and this difference was incorporated into the back-analysis. The measured inclinometer data for wall movement

into the excavation is compared to the back analysis using the original soil strength and stiffness parameters in Table 1. Again, the general shapes of the curves are similar in nature, and the maximum measured wall movement of 48mm (1.9 inches) agrees rather closely with the maximum back-analyzed value of 53mm (2.1 inches), although the elevations at which these maximum values occur are not the same.

The stiffer-than-expected response along the north wall may be partially-explained by the field divergence from plane-strain conditions, but this influence is expected to be less dominant at the north wall, which is longer than the west wall. In addition, the presence of the piled foundation supporting the building to the north may have also contributed to a somewhat stiffer response.

The results of the inclinometer back-analysis indicate that the soil parameters summarized in Table 1 were, in general, quite reasonable.

SUMMARY

This paper has presented the design and construction aspects of a perimeter secant wall for a high-rise real estate development on the west side of Manhattan. The secant wall serves as both excavation support for the basement construction as well as the permanent retaining wall for the new building.

The secant wall was used in lieu of underpinning a two-story building along the eastern site boundary. Since the site is underlain by a thick deposit of very soft organic silt and clay, the potential for excavation-induced ground movements and their impact on the adjacent two-story structure on shallow footings were of the utmost concern.

Design parameters were developed using correlations with the SPT and CPT. The strength of the very soft organic silt and clay layer was characterized by an undrained shear strength (su) ranging from 12 to 24 kPa (250 to 500 psf). The design elastic soil modulus (Es) for this critical layer was established as 250su.

The inclinometer data indicated a maximum lateral movement of 48mm (1.9 inches). In general, relatively good agreement was obtained between these measurements and the results of a back-analysis.

REFERENCES

LANGAN, 2005. Geotechnical Engineering Study – West 19th Street and 11th Avenue, New York, New York, 152 p.

VIELE, 1865. Sanitary and Topographical Map of the City and Island of New York, Ferd. Mayer and Co. Lithographers, New York.

LMW Engineering Group, 2007. Inclinometer Data Report – West 19th Street & Westside Avenue, New York, 4 p.

Figure 13. West Wall Movements

Figure 14. North Wall Movements