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Technical Report Documentation 

1. Report No. 2. Government Accerr~on No. 3. Rec,p*.nt’s Catoloq No. 

FHWA-DP-90-068-003 

4. Title and Subtitl. 5. Report Date 

Pope 

1 

DP-68 - Permanent Ground Anchors 

Volume 2, Field Demonstration 

Project Summaries 

7. Author’s) 

Richard S. Cheney 

April 1990 

6. Pdorm~ng Orqamzotoon Code 

HHO-42 

0. Parformenq Organization Report NO. 

9. Performing Organexation Name and Address 

HHO-30/HHO-40 

10. work unit No. (TRAIS) 

Federal Highway Administration 

400 Seventh Street. SW. 

FJashington, D.C. 20590 

-is s ponsoring Agwwy Name and b.ddr*,* 

HHO-40 

11. Contract or Grant No. 

13. Type of Report and Pwod Covered 

, 

Final Report 

Federal Highway Administration 

400 Seventh Street, SW. 

Washington, D.C. 20590 

IS. Supplwn~ntary NotFHWA: Theodore Ferragut 

1 14. Sponsoring Agency Cod* 

16. Abstract 

Ground anchors, often called tiebacks, are structural elements which receive their 

support in soil or rock and act to retain earth masses and/or applied structural 

loads. The Federal Highway Administration recognized that the use of permanent 

ground anchors in highway cut sections could affect substantial benefits in both 

economy and safety. The specific purpose of the permanent ground anchor demonstration project was to introduce the concept of permanent ground anchor use into 

American construction practice. A manual numbered FHWA-DP-68-1R and titled 

"Permanent Ground Anchors" was prepared and several thousand copies distributed to 

highway engineers. Permanent ground anchors on several projects were monitored to 

validate the concepts addressed in FHWA-DP-68-1R. This report summarizes the 

results of both the field monitoring and the history of the permanent ground anchor 

demonstration project. 

17. Key Words 18. Distribution Statement 

Permanent ground anchors, tiebacks, No restriction - NTIS 

anchors, anchor specifications, anchor 

test projects, anchor acceptance criteria, 

anchor test procedures 

19. Security Clorsif. (of this report) 20. Security Clossif. (of this page) 21. No. of Pago. 22. Pric* 

Form DOT F 1700.7 (8-72) Reproduction of completed page authorized 

REPORT TITLE 

1 Report of Field Performance Permanent Tieback Wall, 

Atlanta, Georgia 

2 Performance of Two Instrumented, Permanent PostGrouted Tiebacks, Baltimore, Maryland 

3 Permanent Tieback Anchors in Cohesive Soil 

Demonstration Site, Seattle, Washington 

4 Performance Monitoring of a Highway Tieback Wall, 

Carroll County, Kentucky 

5 Permanent Ground Anchors for Lateral Support of 

Bridge End Slopes, Anchorage, Alaska 

6 Retention System Monitoring, Lima, Ohio 

ii 

REPORT NO. 1 

REPORT OF FIELD PERFORMANCE 

PERMANENT TIEBACK WALL 

DEMONSTRATION SITE 

I-75 - ATLANTA, GEORGIA 

GEORGIA DEPARTMENT OF TRANSPORTATION 

bY 

LAW/GEOCONSULT INTERNATIONAL 

PROJECT PEI-75-2(41) 

MAY 1985 

SECTION Page 

1 INTRODUCTION........................................4 

2 PROJECT DESCRIPTION ................................ . 

Design Method ................................ ...6 

Construction Method ............................ . 

Subsurface Conditions .......................... . 

INSTRUMENTATION PLAN . . . . . . . . . . . . . . . . . . . . . . . . . . . . . ..g 

Background ..................................... . 

Description .................................... . 

LOAD TESTING PROCEDURE . . . . . . . . . . . . . . . . . . . . . . . . . . ..ll 

TEST RESULTS . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . ...13 

Load Test: Load Displacement . . . . . . . . . . . . . . . . . . ...13 

Load Test: Displacement and Load Distribution 

in Bonded Zone.......................14 

Load Test: Comparison between Load Cell and 

and Hydraulic Jack . . . . . . . . . . . . . . . 19 

Long-term: Load at Anchor Head . . . . . . . . . . . . . . . . ...19 

Long-term: Displacements . . . . . . . . . . . . . . . . . . . . . . . . . 20 

Long-term: Displacements and Distribution 

in Bonded Zone.......................22 

DISCUSSION OF WALL PERFORMANCE . . . . . . . . . . . . . . . . . ...24 

INSTRUMENTATION RECOMMENDATIONS . . . . . . . . . . . . . . . . ...26 

CONCLUSIONS . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . ...30 

Instrumentation Results ......................... ..3 0 

Wall Performance .................................. 30 

Safe Design Loads ............................... ..3 1 

APPENDICES 

Appendix A - Tables and Figures 

Appendix B - Dunnicliff Recommendations for Instrumentation 

l-i 
LIST OF TABLES AND FIGURES CONTAINED 

IN APPENDIX A 

TABLES 

TABLE I - Summary of Instrumentation 

TABLE II - Summary of Acceptance Criteria Creep Testing 

TABLE III - Apparent Modulus of Unbonded Tendon Length Determined 

During Load Tests 

TABLE IV - ,Summary of Anchor Performance 

FIGURES 

FIGURE 1 - 

FIGURE 2-3 - 

FIGURE 4 - 

FIGURE 5 - 

FIGURE 6-15 - 

FIGURE 16 - 

FIGURE 17-18 - 

FIGURE 19-22 - 

FIGURE 23 - 

FIGURE 24 - 

FIGURE 25-30 - 

FIGURE 31-35 - 

FIGURE 36-37 - 

FIGURE 38-40 - 

FIGURE 41-42, - 

. . . . 

FIGURE 43 - 

FIGURE 44 - 

FIGURE 45-46 - 

. 

Plan and Profile Permanent Tieback Wall 

Permanent Tieback Wall Instrumentation 

Schematic of Primary Instrumented Tieback Anchor 

Cross Section of Instrumented Tieback Anchor 

Load Displacement Curve - Load Test 

Creep Curves From Performance Tests 

Comparison of Wire Telltale Versus Rod Telltale 

During Load Test 

Relative Deformation in Bond Zone - Load Test 

Hypothetical Load Transfer Diagrams and Calculated 

Total Elongation of Bonded Zone 

Load Distribution in Bonded Zone Assuming Slippage 

of Wire #l and Short Rod Telltale Anchors 

Comparison of Load Cell and Hydraulic Jack During 

Load Tests 

Long Term Load Variation 

Horizontal Deflection Behind Permanent Tieback wall 

Inclinometer Deflection Versus Time 

Comparison of Inclinometer Top Movement Survey 

. . . . . . . . . . . . . . . . . . Data Versus Inclinometer Data 

Summary of Maximum Horizontal and Vertical Survey 

Measurements 

Horizontal and Vertical Survey Deflections Versus 

Wall Height 

Long Term Relative Deflection in Bonded Zone 

l-ii 
1. INTRODUCTION 

The Georgia Department of Transportation (Georgia DOT) is 

presently conducting an extensive refurbishing program on the 

interstate highways that pass through Atlanta's urban environment. 

An important element of this program is the implementation of many 

new innovations in design and construction in cooperation with the 

Federal Highway Administration (FHYWA) to develop a modern and 

economical interstate system, One such innovation involved the 

design and construction of a permanent tieback retaining wall 

using ground anchors in lieu of a conventional design using a 

cantilevered wall in front of a temporary excavation support 

system. 

The permanent tieback wall design and construction project 

presented in this report was funded through the FHWA's 

Demonstration Projects Program to "promote and accelerate the 

adoption of new research results and innovative planning, 

engineering, and construction practices." The overall objective 

of this demonstration project was to instrument and monitor the 

long-term performance of the wall in order to evaluate the design 

and construction procedures for future Georgia DOT and FHWA-funded 

highway construction projects. This report describes the 

instrumentation plan used to monitor the wall and the ground 

anchors, the testing procedure, 

during 

the test results of data gathered 

the first eight months of monitoring, and presents 

conclusions and recommendations. It is the intent of the Georgia 

DOT and the FHWA to monitor the wall's performance over a three 

year period. 

To accomplish the project objective, 

developed 

a comprehensive work plan was 

to verify the wall design assumptions and construction 

methods and to collect and interpret long-term 

The project scope included the following tasks: 

performance data. 

Task A. Measure deflection of the wall face to an accuracy 

of 0.1 inches. 

Task B. Measure deflection of the wall and ground with 

slope indicators in the unbonded length of the 

tieback and beyond the bonded length. 

. 

Task C. Measure load variations in the tieback tendons 

within the unbonded and bonded lengths at specified 

stations along the wall. 

Task D. Measure other variables which are deemed necessary 

for the evaluation of the wall's performance. 

This demonstration project is one of many that the FHWA has funded 

as part of its efforts to transfer technology into useful 

processes, products and programs. This report presents results of 

l-l 
an innovative permanent tieback retaining wall used for the first 

time by the Georgia DOT. The dissemination of the results will 

assist other State highway agencies to determine the applicability 

of permanent tieback walls to their highway construction needs. 

2. PROJECT DESCRIPTION 

This demonstration project is included as part of Georgia DOT's 

Project P.E. I-75-02 (41) in Fulton County. This construction 

project involved the widening of I-75/85 in Atlanta near Fifth 

Street and in close proximity to the Penthouse Motel and the Coca 

Cola Bottling Company. The widening required the construction of 

a retaining,wall as shown in Figure 1. The permanent tieback 

retaining wall section is approximately 500 feet long and its' 

height ranges from seven to 30 feet. This site was selected to 

demonstrate the permanent tieback design because of the right-ofway limitations and potential effects of construction on the 

adjacent structures which limited the type of retaining wall that 

could be built at this location. 

Desisn Method 

The original design, provided by Georgia DOT and the FHWA, 

included both a cast-in-place concrete cantilever wall and a 

permanent anchor wall as bid alternates. The general contractor, 

S. J. Groves, chose the permanent anchor solution and submitted a 

revised wall designed by Nicholson Construction Company and their 

subcontractor, Chastain and Tindel, Inc. This revised permanent 

anchor wall design was reviewed and accepted by Georgia DOT and 

the FHWA. The design loading on the wall included active earth 

pressure, and impact and surcharge loads, which were included due 

to the close proximity of traffic on Williams Street. These 

loads were to be resisted by the pile embedment (passive earth 

pressure) and the tieback loads. Active earth pressure loading 

was assumed,to be a uniform distribution acting on the soldier 

piles and wood lagging above excavation levels. This distribution 

was assumed to be 65 percent of the computed active case 

hydrostatic load at the excavation level, as described by Terzaghi 

and Peck (1967). Hydrostatic earth pressure distributions were 

assumed to act on the soldier piles below excavation levels. The 

active component of the hydrostatic load was assumed to act on a 

2.5 foot wide pile (effective width of concrete filled drilled 

hole). Passive earth pressure resistance was assumed in front of 

the piles below excavation levels. In addition to the 2.5 foot 

wide pile, as allowed by the specifications, an effective pile 

width factor of 3 was used. This factor allows the distribution 

of passive earth pressure to act on an effective pile width (7.5 

feet) to account for a soil arching effect in front of the piles. 

While no direct factor of safety was applied to the passive earth 

pressure coefficient, a safety margin was included in the design 

by assuming only one-half of the actual pile embedment in the 

l-2 
calculations. No water pressure loading was assumed to act on the 

retaining wall since groundwater levels were deeper than 

excavation levels. Also, drainage fabric was to be installed 

behind the concrete facing to carry away any groundwater that 

might collect behind the wall. 

Tieback loads were calculated using a structural computer program 

and were checked by hand calculations. Both methods consisted of 

summing forces and moments and checking for equilibrium. One 

level of tiebacks was included in the lower portion of the wall, 

while the remaining portion of the wall included two levels. 

Design tieback loads ranged from 57 kips (in the lowest portion of 

the wall) to 138 kips. In cases where the design load in a 

tieback was not achieved during field proof testing, the 

requirement for an additional tieback was also calculated by 

summing moments and forces. 

The soldier piles were designed as simply supported steel beams 

where two or more tiebacks were installed on the soldier pile. 

Both final configuration and critical construction sequence 

loadings were examined. The latter calculations were made for 

intermediate excavation levels to check for maximum moments in the 

piles occurring during construction rather than at the deepest 

excavation levels. 

Construction Method 

The method of construction used in this tieback wall involved 

several components and sequences. The wall consisted of soldier 

piles, timber lagging, ground anchors and concrete face panels. 

First, 30 inch diameter holes were augered from the ground surface 

to design depths ranging from seven to nine feet below the 

proposed bottom of the wall. Soldier piles were inserted into the 

holes. The piles were located on eight or nine foot centers 

approximately 15 inches behind the proposed wall face. 

The structural members used for the soldier piles were double 

beams (ranging from W12 x 26 to W18 x 46 sections) joined with 

steel plates. Concrete (3000 psi strength) was placed around the 

piles from the bottom of the hole to the proposed ground line. 

Lean concrete was placed above the proposed ground line. As earth 

in front of the piles was excavated, the lagging was installed 

between the piles bearing against the steel beam flanges. At 

designed locations and angles, approximately 6 inch diameter holes 

were drilled between the two joined beams through the lean 

concrete and into the soil for placement of the ground anchors. 

Casing was placed as the hole was advanced to allow insertion of 

the anchor. The ground anchors were stranded tendons 

approximately 50 feet long placed at angles from 20 to 30 degrees 

from the horizontal. Each tendon consisted of from two to five 

strands with each strand made up of seven wires. The majority of 

tendons in the higher portions of the wall had either four or five 

strands. The five strand tendon had an approximate diameter of 

1-3 
0.6 inches and total cross-sectional area of 1.075 square inches. 

In the unbonded portion of the tendon, a grease filled plastic 

sheath covered each strand. A corrugated plastic sheath was 

fitted around the entire tendon in both the unbonded and bonded 

zone. Grout was injected through the casing as the casing was 

withdrawn from the hole. The grout eventu,ally filled boththe 

annular space between the tendon and the corrugated tube, and the 

corrugated tube and the ground. Grout pressures of approximately 

70 pounds per square inch were utilized. Grout pressures were 

measured with gages located at both the drill head and grout pump 

to ensure that the pressure was maintained. 

After curing of the grout, the anchors were load-tested and post 

tensioned to a predetermined load. The load test procedure is 

described in detail in Section 4. In some cases, an anchor would 

not hold the predetermined design load. In these cases, the 

anchor was assigned a "safe" load based on the load test results. 

The designer would then assess whether this reduced load capacity 

was sufficient. If not, an additional anchor was installed. 

Additional anchors were required at the northern instrumentation 

section (Piles 65, 66, and 67). 

The process of lagging installation and anchor installation was 

repeated until the excavation in front of the wall was complete. 

The final sequence was to install strips of drainage fabric along 

the lagging and cast concrete face panels over the piles, lagging 

and tieback retaining system. The concrete face panels were 

attached to the soldier piles by a series of studs welded to the 

soldier piles and embedded in the cast-in-place concrete. 

Subsurface Conditions 

Subsurface conditions at the site were determined by borings 

performed by the Georgia DOT. The over-burdened soils were 

described as "medium dense to dense yellowish-brown and pinkish 

micaceous sandy silts." The micaceous nature of the site soils 

was of interest because of its potential effect on anchor capacity 

and particularly long-term creep. The residual soils in the 

Piedmont geologic province are typically micaceous due to their 

derivation from parent geneses and sects. Mica content has been 

estimated (Sowers, 1963) in the Southern Piedmont area 

typically range from 5 to 25 percent. Quantification of soil rnikz 

content is not commonly performed; therefore, it was not included 

by Georgia DOT as part of their investigation. 

Standard Penetration Test resistances ranged from 11 blows per 

foot to 60 blows with no penetration. These values generally 

increased with depth. However, harder and softer layers were 

frequently encountered in interbedded fashion. Groundwater was 

encountered approximately 40 feet below the ground surface. The 

depth to hard rock varies from 45 to 65 feet. 

During construction, thin black, iron-manganese coated zones were 

observed. These zones are common in the Southern Piedmont and are 

l-4 
locally termed "slickensides," but are actually unrelated to 

faulting (St. John, et al, 1969). 

3. INSTRUMENTATION PLAN 

Backaround 

The instrumentation of the permanent tieback wall consists of 

several mon‘itoring devices. The instrumentation plan was 

recommended by Mr. John Dunncliff and is appended to this report. 

Table I briefly presents all monitoring devices, including type and source of instruments and their primary purpose within the 

program. A discussion of the instrument plan is contained in the 

following paragraphs. 

A significant change resulted early in the instrumentation program 

when stranded tendons were selected over the initially planned bar 

anchors. Strain gages were initially planned to measure strain at 

designated locations within the bonded zone and thus allow 

calculation of load transfer. However, it was not deemed feasible 

to mount strain gages to the tendon strands. Thus, wire 

telltales, with a custom-fitted fixation to the strand, were used 

to measure deformation within the bonded zone. 

Description 

Two monitoring stations spaced approximately 135 feet apart along 

the wall were selected to monitor the wall performance in higher 

portions of the wall (Figures 2 and 3). At each monitoring 

station, two production anchors (upper and lower) on a single pile 

line were designated for "primary" instrumentation. The 

production anchor on the piles on both sides of the "primary" 

anchor was designated for "secondary" instrumentation. The 

instrumentation for two secondary anchors was destroyed during 

construction (discussed in Section 7). Thus, a total of four 

primary and six secondary anchors were instrumented. 

Instrumentation for the primary instrumented anchors included: a 

permanent load cell; short and long rod telltales; and five wire 

telltales. This arrangement is shown in Figure 4 . The five wire 

telltales were fixed approximately 7.5 feet apart from the top to 

the bottom of the bonded zone (Instrument Positions 1 through 5). 

The short rod telltale was fixed at the top of the bond zone 

(Instrument Position 1). The long rod telltale was fixed at the 

bottom of the bond zone (Instrument Position 5). Rod and wire 

telltales were attached to the same tendon strand. A basic 

assumption in application of the wire and rod telltales was that 

all strands within the anchors behaved similarly and that 

instrumentation of one strand would predict the behavior of the 

entire anchor. 

l-5 
The secondary instrumented anchors had only one rod telltale fixed 

to one of the five strands at Position 1. The load cell mounted 

at the anchor head during load testing was not left in place 

permanently. 

In addition to the anchor instrumentation, each monitoring station 

included three inclinometers. One inclinometer was placed in a 

pipe which was welded to the soldier pile for the primary anchors 

(Piles 51 and 68). The other two inclinometers were located 

approximately 10 and 50 feet behind the wall. 

Extensive horizontal and vertical survey data all along the 

permanent tieback wall was provided by the Georgia DOT. Survey 

points were established on the wall top and face. Ground survey 

points were also established at several distances (10 and 30 feet) 

behind the wall. Points 50 feet behind the wall were included for 

a limited portion of the wall. Movement of the inclinometer tops 

was also measured. Horizontal survey was performed by electronic 

distance measurement from points across the existing interstate. 

Vertical survey was performed by routine level measurements. 

Each instrument was given a code for identification. This code is 

presented below. 

w - XY - z 

where W = Instrument Type 

RT = Rod telltale 

WT = Wire telltale 

LC = Load cell 

SI = Slope inclinometer 

x = Soldier Pile Number 

See Figures 2 and 3 

1 = Anchor Locations 

U= Upper Row or L = Lower Row 

z = Instrument Position Desiqnation 

See Figure 4 

4. LOAD TESTING PROCEDURE 

All project anchors were subjected to a load test during 

installation. The load tests were either a "proof test" or a 

l-6 
"performance test". All anchors were to be loaded to 150 percent 

of their design load. Incremental cycling of load was part of the 

performance load testing. Anchors were loaded with a center-pull 

hydraulic jack with an electric pump. Displacement of the loading 

head was measured with a dial gage attached to a tripod set on the 

ground. 

The acceptance criteria for the anchors was based on the movement 

of the anchor head (measured by the dial gage) during load testing 

and creep movement at the maximum test load. An anchor was 

acceptable if: 

A. The total elastic movement obtained exceeded 80 percent 

of the theoretical elastic elongation of the free length 

and was less than the theoretical elastic elongation of 

the free length plus 50 percent of the bond length. 

B. The creep movement did not exceed 0.080 inches during a 

single log-time cycle (five to 50 minutes for 

performance testing, and one-half to five minutes for 

proof testing) regardless of tendon length and load. 

Anchors which met the above acceptance criteria were unloaded and 

locked-off at 115 percent of design load. The additional 15 

percent of load was included to account for load losses due to 

lock-off (wedge setting) and potential long-term creep losses. 

Anchors which did not meet the above criteria were redesigned. 

This typically involved assuming a new design load equal to 67 

percent of the maximum safe load determined during the load test. 

The anchor was unloaded and locked-off at 115 percent of the "new" 

design load. 

Both anchors which met the original criteria and redesigned 

anchors were actually unloaded to below the design load and then 

reloaded to the lock-off load. This is because the jack pressure 

(which determines load) is considered to be more accurate in 

loading than in unloading. Thus, by loading up to the lock-off 

load, a more accurate reading is obtained. 

Several d-ifferences in the load test procedures were utilized for 

the instrumented anchors (primary and secondary). A 150-ton and a 

300-ton capacity hydraulic jack was utilized for loading the 

secondary anchors and primary anchors, respectively. The primary 

anchors required a larger jack because the center hole in the 150- 

ton jack was not large enough to accommodate the instrumentation 

and the anchor. In addition to the jack load reading, a load cell 

was used to monitor the anchor loads. For this reason, it was not 

necessary to unload and reload the anchors to obtain an accurate 

lock-off load. A different loading sequence occurred at four 

anchor locations. At anchor locations 52-upper, 67-upper and 67- 

lower the anchors were unloaded to the seating load (alignment 

load) and then loaded back to the lock-off load. At anchor 

location 66-lower, the load was cycled (loaded and unloaded) at 

l-7 
approximately 50 percent of design load and cycled again at 150 

percent of design load. Some details of the load testing at each 

of the instrumented anchors are shown on Figures 6 through 15 

(Load Displacement Curve - Load Test). 

5. TEST RESULTS 

Test results presented and discussed within the following section 

are divided into those obtained during load testing of anchors and 

those obtained during the long-term (approximately eight months) 

monitoring program. 

Load Test: ,Load-Displacement 

Loads applied during load testing of both primary and secondary 

anchors were measured with both the jack pressure gage and load 

cell. However, the load data presented within this section were 

measured with the load cell because it is generally considered to 

be more accurate. Displacements of the anchor head were measured 

with a dial gage which read displacement independent of anchor 

movement. Plots of load versus head displacement are shown in 

Figures 6 through 15. The acceptance criteria (A) discussed in 

the Section 4 are also shown on these figures. The free and bond 

lengths used for determination of the acceptance criteria are 

shown on Figures 6 through 15. Some results of the load tests 

have been included in Table IV, Summary of Anchor Performance. 

Displacement of the long and short rod telltales and wire 

telltales. #l and #5 for the primary anchors are also shown on 

Figures 6 through 9. The long rod telltale and wire telltale #5 

measure displacement of the pulling head with respect to the 

anchor bottom. As may be seen only minor differences exist 

between the dial gage and long rod telltale readings for primary 

Anchors 51-upper, 51-lower, and 66-lower. This indicates that the 

bottoms of these anchors showed no or very little movement during 

the load test. Anchor 66-upper (Figure 8) indicates a difference 

of approximately 0.1 inches between the dial gage and the long rod 

or wire #5 telltale. Thus, a displacement of this amount of the 

anchor bottom may have resulted during load testing. 

The modulus of elasticity of the unbonded length can be determined 

during the load test for the primary and secondary anchors by 

comparison of load versus elongation of the unbonded length. This 

elongation is measured with the short rod telltale, or wire #l, 

both of which are anchored at the interface between the unbonded 

and bonded zones. Modulus values determined in this manner are 

presented in Table III. The ,reported manufacturer's modulus was 

28.5 x lo6 pounds per square ,inch. As may be seen in that table, 

the calculated modulus values during loading are considerably less 

than the manufacturer's modulus. A potential explanation for this 

difference in slippage of the rod and wire telltale anchors 

l-8 
relative to the tendon, causing the telltales to overestimate 

elongation. The slippage could be caused by the tendon 

out" of the grout 

'*breaking 

with the telltale anchors restrained by the 

grout. The \possibility of slippage cannot be discounted, despite 

a soft material which was placed immediately in front of the 

telltale anchors (around the tubes containing the rods or 

designed to alleviate this problem. 

wires) 

As discussed later, an 

evaluation of load transfer data tends to support 

of slippage of these telltale anchors. 

the hypothesis 

As discussed previously, five and 50 minute creep tests were 

performed at the maximum test load for the secondary and primary 

anchors, respectively. 

in Table II. 

The results of these tests are presented 

We have also utilized the creep test data to make a preliminary 

assessment of creep potential. For this assessment, we have used 

a procedure presented in the report "Permanent Ground Anchors: 

Soletanche Design Criteria," dated September 1982 (Report No. 

FHWA/RD-81/150). This procedure involves the performance of 

several "design load tests" prior to the installation of 

production anchors. In these tests, the anchor load is held for 

one hour at different increasing load levels. These "design load 

tests" were not performed as part of either the design or 

construction of the permanent tieback wall. However, we have 

utilized the, results from 50 minute creep tests performed both for 

the primary instrumentation anchors 

testing for selected production anchors. 

and as part of performance 

This data is plotted 

(Figure 16) in a manner similar to that suggested in the 

Soletanche Report. The data appears to generally suggest that the 

"critical creep tension" was not reached during these tests. 

Load Test: Displacement and Load Distribution in Bonded Zone 

Displacement along the bonded anchor zones of the four primary 

anchors was measured using a series of five wire telltales. The 

spacing between the wire telltales, 

approximately 7.5 feet. 

as shown in Figure 4, was 

bonded 

Displacement at the sop and bottom of the 

zone was also measured with the short and 

telltales which were attached at the 

long rod 

strand as 

same place on the tendon 

the wire telltales #l and #5, 

comparison of the short rod versus 

respectively. 

versus wire #5, 

wire #l, and the long rot 

is shown in Figures 17 and 18 (Appendix A) for the 

four primary anchors. Because the rod and wire telltale are fixed 

at the same location 

approximately 

on the tendon strand, they should measure 

equal deflection (that is, approximately zero 

relative deflection on Figures 17 and 18). 

agreement (less than 0.05 inch relative deflection) 

Relatively good 

was obtained 

for Anchors 51-upper, 66-upper, 

rod, only). 

and 51-lower (wire #l versus short 

Lesser agreement was obtained for Anchor 86-lower 

(0.05 to 0.10 inch relative deflection) and Anchor 51-lower, wire 

#5 versus long rod (0.10 to 0.15 inch relative deflection). For 

both Anchors 51-lower and 66-lower we have also shown relative 

l-9 
deflection of the long rod versus dial gage, which should be 

approximately equal if no deflection occurs at the bottom of the 

bonded zone. Good agreement between the dial gage and long rod 

was obtained for Anchor 51-lower, thus indicating that wire #5 

overestimated deflection. In Anchor 88-lower, the dial gage was 

greater than the long rod which could indicate either displacement 

at the bottom of the bonded zone or relative instrument accuracy. 

Wire #5 deflection exceeded dial gage deflection, which probably 

indicates. that wire #5 overestimated deflection. 

The displacement measured by each wire telltale with respect to 

wire telltale #l (relative deformation) during load testing is 

shown on Figures 19 through 22. The difference between wire 

telltale curves at a given load represents the relative 

deformation between points. Example: the difference between 

curves 2-l j and 3-l represents the relative displacement between 

points 2 and 3). Positive relative deformations between any two 

points (2-1, 3-1, 3-2, etc.) represents elongation of that segment 

of the tendon. The displacement of the long rod telltale with 

respect to the short rod telltale is also shown on these figures. 

The relative deformations within the bonded zone must meet two 

criteria which are based on two closely related assumptions. 

These criteria and their related assumptions are: 

A. Relative deformation (elongation) should increase with 

increased distance between points (the 3-l difference 

should be greater than the 2-l difference, etc.) 

assuming that the tendon is in tension. The wire 5-l 

difference long rod-short difference 

(whichever is izre %?rect) should form tEzdupper bound 

of the plots shown on Figures 19 through 22. 

B. The elongation of adjacent equal length segments must 

decrease moving down the tendon. That is, the wire 2-l 

difference must be greater than the wire 3-2 difference 

(3-2 = 3-l minus 2-l), etc. This assumes that tension 

in the tendon decreases with increased distance from the 

top of the bond zone. 

The relative deformations presented in Figures 19 through 22 may 

be examined utilizing these two criteria. Anchor 51-lower, with 

the displacements measured by wire #3, #4, and #5 all exceeding 

the long rod displacement, will be excluded from this 

consideration. Examination based on data for the remaining three 

anchors (Figures 19, 21, and 22) yields the following general 

points: 

A. The wire 2-l difference is less than the wire 3-l 

difference (Criterion I) in all three cases (including 

Anchor 66-upper for which the wire 2-l difference was 

negative). However, the wire 2-l difference is less 

than the wire 3-2 difference (3-2 = 3-l minus 2-l) in 

all three cases, which is contrary to Criterion II. 

l-10 
B In two of three anchors (88-upper and 66-lower) the wire 

3-l difference is less than the wire 4-l difference 

(Criterion I). In both cases, the wire 4-3 difference 

is less than the wire 4-3 difference (Criterion II). 

c. The wire 4-l difference is less than the wire 5-l 

difference in Anchor 

wire 3-l difference) 

51-upper (and also less than the 

and close to the wire 5-l 

difference in Anchors 66-upper and 66-lower. In 

general, deformations in this portion of the bonded zone 

appear to be in the range of the data accuracy. 

it appears difficult to 

Thus, 

resolve deformations in this 

portion of the bonded zone. 

In summary, the wire 2-l difference appears too small in all 

cases, while the validity of the remaining data appears mixed. A 

potential reason for this apparent underestimation of the 2-l 

relative deformation (elongation) is slippage of the wire #l (and 

short rod) telltale anchorage. 

slippage could 

As discussed previously, this 

cause an overestimation of elongation of the 

unbonded length and, thus, an underestimation of elongation of the 

bonded zone (particularly the upper 7.5 feet of the bonded 

when wire #l is compared to wire #2). 

zone, 

anchors, 

Slippage of the telltale 

caused by the grout restraining free movement of the 

telltale anchor along with the tendon, appears more likely at 

Point #l than at the other telltale anchorage points. This is 

because higher tensions and displacements occur in this portion of 

the tendon than in deeper portions of the bonded zone. 

tensions 

The higher 

and displacements increase the possibility that the bond 

between the tendon and the grout could be broken and thus that 

relative displacement between the two could occur. 

the grout within the unbonded zone (free length) is 

In addition, 

not stressed 

by the tendon and thus acts to restrain the outward movement of 

grout in the upper portion of the bonded zone. 

The amount of possible slippage of the wire #l and short rod 

telltale anchors can be estimated by making the assumption that 

telltale anchor slippage is the only cause of difference between 

the apparently underestimated modulus values determined during the 

load test and the manufacturer's modulus. the difference 

between 

Thus, 

the elongation compatible with the manufacturer's modulus 

and the elongation measured by wire #l (or the short rod) could 

approximately equal the amount of slippage. The amount of 

estimated slippage increases with load and is 

approximately 

increasing 

0.3 inches at maximum loads. If the wire #l and 

short rod readings are corrected for the estimated 

the wire differences (2-1, 

slippage, all 

3-1, 4-1, 5-1, and the long rod-short 

rod) plotted on Figures 19 through 22 would be moved to the 

with respect to the ordinate. 

right 

The purpose of deflection measurements within the bonded zone was 

to provide an indirect calculation of load transfer. 

of slippage of 

The question 

the wire #l and short rod telltale anchor 

l-11 
significantly effects the results of this calculation. This 

effect will be further discussed below. First, however, we will 

discuss the manner in which load transfer is estimated from 

relative deflections within the bonded zone. 

Elongation of a segment of the tendon is a function of strains 

along that segment length. 

of 

The elongation equals the 

strain 

integration 

plot). 

versus length (the area under the strain versus length 

Thus, 

within 

elongation is directly related to the average strain 

that segment. Stress (and load) are directly related to 

strain at any point along the tendon (Hooke's Law). Thus, the 

shape of load versus length plot must be directly related to the 

shape of the strain versus length plot, and average load directly 

related to, average strain. Further, the average load within a 

segment, as 'a measure of the area under the load distribution 

curve, is directly related to the elongation of the segment, and 

can be calculated from that elongation. The simplifying 

assumption that strains and loads change linearly within a segment 

allows positioning the calculated average load at the mid-point of 

the segment. The change in load from one point in the tendon to 

another point is the load transfer between those points. 

As discussed previously the question of slippage of the wire #l 

and short rod telltale anchors significant affects our calculation 

of load transfer. Thus, the two possible outcomes of the question 

lead to two separate lines of reasoning, which are developed 

separately below: 

A. Assumption of No Telltale Anchor Slippaqe: If we assume 

that slippage of the wire #l and short rod telltale 

anchors has not occurred, then the wire 2-1 differences 

become too small compared to the wire 3-2 differences 

(Criterion II). Thus, both these wire differences 

appear in question and it appears impossible to utilize 

them to calculate average loads and load transfer in the 

upper 15 feet of the bonded zone. However, if slippage 

has not occurred, then a good estimate of overall bond 

zone elongation appears possible utilizing the wire 5-l 

or long rod-short rod difference. This data is 

supported by the previously discussed agreement between 

the short rod versus wire #l, and the long rod versus 

wire #5. A significant result of the previous 

discussion on calculation of load transfer is that 

elongation of the entire bonded zone (or any segment of 

the bonded zone) is directly proportional to the area 

under the load distribution curve. Thus, although it is 

not possible to calculate average loads within segments 

of the bonded zone, it may be possible to derive general 

trends in the load distribution. Anchors Sl-upper and 

66-lower both showed approximate 0.6 inch bonded zone 

elongation at 150 kips above the seating load. Thus, 

Part A of Figure 23 shows three hypothetical load 

distribution diagrams with elongation of the entire 

1-12 
bonded zone equal to 0.6 inches. Examination of these 

diagrams shows that significant load transfer must occur 

within the first two segments of the bonded zone to 

result in 0.6 inches of elongation. Lesser load 

transfer rates in this portion of the bonded zone would 

result in increased area under the curve and thus 

increased bonded zone elongation. 

B. Assumption of Telltale Anchor Slippaae: If we assume 

that slippage of the wire #l and short rod telltales has 

occurred, 

through 22 

then the wire difference curves on Figures 19 

are shifted to the right, as discussed 

previously. As stated, we may calculate the amount of 

slippage based on the manufacturer's modulus and use it 

to correct the wire #l data, This increases the wire 2- 

1 difference, but does not affect the relative 

differences between the other curves (the wire 3-2, 4-3, 

and 5-4 differences). The overall elongation of the 

bond zone (the wire 5-l or long rod-short rod 

difference) is, of course, also increased. The maximum 

slippage, as previously discussed, is estimated to be 

approximately 0.3 inches. Adding this amount to the 0.6 

inches discussed previously for Anchors 51-upper and 66- 

lower, yields a total bond zone elongation of 

approximately 0.9 inches. Part B of Figure 23 shows 

hypothetical load distribution diagrams with elongation 

of the entire bonded zone equal to 0.9 inches. In 

addition, we may use the adjusted wire 2-l difference 

and the other wire differences to calculate average 

segmental loads and thus derive a load distribution plot 

which is based on the slippage assumption. This has 

been performed for Anchors 66-upper and 66-lower. After 

adjustment of the data from all four anchors for assumed 

slippage, we judge the data from these two anchors to be 

more reasonable than the other two primary anchors. 

Load distribution plots for these two anchors are shown 

on Figure 24. Load distribution for selected loading 

increments are shown. At each load increment we have 

plotted calculated average load within bonded zone 

segments. We have also shown smoothed load 

distribution curve which meets thz criteria, stated 

previously, 

related 

that area under the curve must be directly 

to elongation of the entire bonded zone. These 

curves intersect the ordinate of the plots at the load 

cell reading, which is assumed to equal load at the top 

of'the bonded zone. The calculated average 

loads 

segmental 

and the smoothed load distribution curve are both 

based on deflection data which begins at the 

load. 

seating 

Thus, these loads must be added to assumed load 

distribution at the seating load. This assumed 

distribution is shown on the plots. 

In summary, our estimate of load distribution within the 

bonded zone is significantly affected by the question of 

1-13 
slippage of the wire #l and short rod telltale anchors. 

The assumption of no slippage results in our inability 

to utilize the wire differences to estimate loads, and 

results in relatively high load transfers within the 

upper portions of the bonded zone (based on total 

elongation of the bonded zone). The assumption of 

anchor slippage is based on the low modulus values 

calculated for the unbonded zone during the load test. 

Application of the estimated slippage makes the wire 

differences appear more reasonable and allows 

calculation of average segmental loads. Load transfer 

b&sed on the anchor slippage assumption appear more 

reasonable than that based on the no slippage 

assumption. Thus, there appears to exist several points 

supporting the slippage assumption. However, we do not 

feel that it is possible to decide conclusively between 

these two assumptions and associated results. Future 

projects might provide insight into this question, 

Load Test: Comparison between Load Cell and Hvdraulic Jack 

Figures 25 through 30 show jack load reading divided by load cell 

reading versus load cell reading during load testing for primary 

and secondary anchors. The load cell reading was chosen for the 

abscissa because it is assumed to be more accurate that the jack 

load reading. As may be seen the jack load reading in almost all 

cases is greater than or equal to the load cell during loading and 

less than the load cell during unloading. The load as measured by 

the jack has an apparent accuracy of + 5 to 10 percent. This 

agrees with most published data which generally cite off-center 

loading and end effects as probable cause of inaccuracy. 

Lonq term: Load at Anchor Head 

Long-term variation in load at the anchor head was measured on the 

primary anchors with load cells and short rod telltales (telltale 

load cells), and on the secondary anchors with only short rod 

telltales. 

The functioning of the short rod telltale as a load indicator is 

as follows. (A sketch and description are also contained in 

Figure 3 of the appended instrumentation plan. The short rod 

telltale measures displacements between the top of the bonded zone 

and the anchor head. Thus, the short rod telltale measures 

changes in length of the unbonded anchor strand. These changes in 

length can be converted into load changes (via Hooke's Law) within 

the unbonded zone. It is assumed that no load transfer occurs 

within this unbonded zone. Thus, these load changes are also 

indicative of load changes at the anchor head. Because the wire 

telltale -#l also measures changes in length of the unbonded zone, 

it can be used as a telltale load cell in a similar manner. The 

unbonded length in the final lock-off position (the length 

utilized in the telltale load cell calculations) is less than the 

1-14 
unbonded length during the load test. This is because a portion 

of tendon was cut off during the lock-off. This reduction in 

length was typically 4 to 5 feet. 

Utilization of Hooke's Law, as discussed above, requires 

assumption of a modulus of elasticity 

w l As discussed 

previously, modulus values of the unbonded length determined 

during load testing were significantly less than the reported 

manufacturer's modulus and may have been caused by slippage of 

wire #l and short rod telltale anchors. For this reason, it was 

decided to use the manufacturer's modulus. 

Long-term load variation at the anchor head is shown for the 

primary anchors on Figures 31 and 32. Load change measured by the 

load cell and calculated by both the rod and wire telltale load 

cells are shown. As may be seen, there is good agreement in three 

of four primary anchors between the loads changes measured with 

the load cell and those calculated from the rod telltale. This 

agreement tends to support the decision to use the manufacturer's 

modulus in the telltale load cell calculations. The difference 

between the load indicators in Anchor 51-lower cannot be 

explained. Based on the comparisons, it appears that the short 

rod telltale is a relatively reliable indicator of long-term load 

change. Figures 33 through 35 show load variations in the six 

secondary anchors as measured with the rod telltale load cell. 

Measurement of long-term load change for both primary and 

secondary anchors was begun after setting wedges to lock-off the 

anchor into final configuration. 

Approximately eight months of load variation data has been 

collected for all ten instrumented anchors. Seven of the anchors 

have shown approximately 5 kips or less variation from their lockoff condition (Anchors 50-upper, 51-upper, 51-lower, 52-upper 

per, 65-upper, 65-lower, and 66-upper). Of the remaining three 

anchors, two have shown increase in load of approximately 10 kips 

(Anchors 66-lower and 67-lower) while Anchor 67-upper has shown a 

decrease in load of approximately 15 kips. The load changes after 

eight months have been added to the seating loads, and are shown 

in Table IV. We have also attempted to estimate the long-term 

load loss (creep) in anchors. These results are also shown in 

Table IV. These estimated creep values should be considered 

preliminary and should be reevaluated when additional long-term 

data is available. 

Lonu-term: Displacements 

Three inclinometers (SI) are located behind each instrumented 

station (Piles 50-52 and 65-67). The Number 1 inclinometers were 

installed in a steel pipe welded to Piles 51 and 66. The Number 2 

and 3 inclinometers are located in the soil approximately 10 and 

50 feet behind the retaining wall, respectively. The horizontal 

inclinometer deflections observed at 

construction are shown in Figures 36 and 37. 

different stages of 

1-15 
The inclinometer casings were extended approximately 10 feet below 

the bottoms of the soldier piles as shown in Figures 36 and 37. 

As may be seen by examination of 

those figures, 

the inclinometer profiles on 

casings. 

fixity was not exactly achieved at the base of the 

The amount of base deviation was small and the effect on 

the overall readings appears negligible. This 

confirmed by horizontal survey of the 

is generally 

inclinometer 

Comparisons 

casing tops. 

in Figures 

of survey data versus inclinometer top data are shown 

41 and 42. 

determine 

The procedure of "check sums" was used to 

inclinometer accuracy. This confirmed that the 

inclinometers were providing reliable data. 

The maximum',horizontal retaining wall movement measured 

inclinometers 

by the 

was about 3/4 inches toward the excavation. The 

magnitude of the horizontal movements 10 feet behind the wall 

(Number 2 inclinometers) are similar 

wall; however, 

to that of the retaining 

the vertical distribution of the movement is 

somewhat different. Somewhat greater deflection occurred in 

Inclinometer SI-66-2 than in Inclinometer SI-66-1. The deflection 

of Inclinometer SI-51-3 (approximately 50 feet behind the wall) 

was not significant. However, 

40 feet behind the wall) 

Inclinometer Sl-66-3 (approximately 

measured approximately 0.4 inches 

deflection. 

Figures 38 through 40 show inclinometer deflection at given depths 

versus time. As may be seen, movement continued to occur after 

the excavation was complete but appears to be 

time, especially over the last three months. 

slowing down with 

Extensive horizontal and vertical survey data was provided by the 

Georgia DOT. Survey points were established on the wall 

face six feet below the top of top and 

the wall. 

established on the wall 

All survey points 

top useful long-term 

were destroyed prior to yielding 

information. Ground survey points were also 

established,at several distances (10, 30 and 50 feet) behind the 

wall. As 'discussed above, 

also measured. 

movement of the inclinometer tops was 

Maximum horizontal and vertical deflections are 

shown in plan view on Figure 43. Maximum horizontal and vertical 

deflections as a function of wall height are plotted in Figure 44. 

Survey points on the wall face generally indicated an inward 

movement during the time anchors were stressed. This inward 

movement ranged from approximately 0.03 to 0.07 feet (0.4 to 0.8 

inches). Outward movement was referenced from 

maximum inward movement and 

the point of 

was a maximum of 

inches). 

0.10 feet (1.2 

Ground survey points 10 feet and 30 feet behind the wall 

face showed a maximum outward horizontal deflection of 0.11 feet 

(1.3 inches) and 0.09 feet (0.8 inches), respectively. 

seen in Figure 44, 

As may be 

a good correlation apparently exists between 

wall height and surveyed outward deflections of both the piles 

(measured from the point of maximum inward deflection) and ground 

points. Examination of the survey data appears to indicate that 

wall movement has stabilized. 

1-16 
Downward deflections of the surveyed pile monitor points indicated 

a downward maximum deflection of 

inch). 

approximately 0.08 feet (1.0 

As may be seen in Figure 44 these deflections were erratic 

with respect to wall height. The pile settlement was probably 

caused by the downward component of tieback or 

loads. 

earth pressure 

Ground survey points 10 feet and 30 feet behind the wall 

showed a maximum settlement of 0.07 feet (0.8 inch) and 0.02 feet 

(0.2 inch), respectively. As may be seen in Figure 44 a good 

correlation apparently exists between 

ground settlements. 

wall height and surveyed 

Surveyed vertical ground deflections averaged 

approximately 50 percent of horizontal deflections. 

Maximum horizontal wall movements measured by the inclinometers 

and by survey were approximately 0.8 inches and 1.2 inches, 

respectively. These correspond to approximately 0.2 percent and 

0.3 percent of wall height, respectively. Goldberg, Jaworski and 

Gordon (1976) reported a range of horizontal movements of 0.1 

percent and 0.8 percent of wall height for tieback walls in sand 

and gravel. Experience with standard cantilever retaining walls 

in the Atlanta area indicates that development of the active earth 

pressure case requires a horizontal wall movement of approximately 

0.3 to 0.6 percent of wall height. 

the permanent tieback wall are in the 

In our opinion, movements of 

associated 

range of 

with 

those generally 

the active earth pressure condition for soils in 

the Atlanta.area. 

One purpose of the survey data was to determine whether or not the 

instrumented sections were typical of the rest of the wall. 

Examination of Figure 43 and 44 appears to indicate that they were 

typical of the higher portion of the wall. 

Lonq-term: Displacements and Load Distribution in Bonded Zone 

The wire #5 and long rod telltales measure relative deformations 

between the anchor head and the bottom of the bonded zone 

the entire anchor). 

(i.e. 

Examination of the long term data indicates 

elongations of the anchor ranging from 0.05 to 0.20 inches. This 

is compared to outward movements of the wall and anchor head, of 

approximately 3/4 inch. Thus, as the wall deformed the 

entire anchor 

outward, 

(including the bonded zone) also moved outward 

approximately l/2 to 3/4 inch. Inclinometer SI-66-3 was located 

near the bottom of the anchors (see 

approximately 0.4 inches of deflection. 

Figure 37) and measured 

Thus, movement of the 

bottom of the anchors may be related to overall movement of the 

soil mass rather than displacement of the bonded zone in relation 

to the surrounding soil. 

Overall elongation of the entire anchor, 

#5 and long rod telltale, 

as measured by the wire 

consists of unbonded length deformation 

(indicative of load in the anchor) and bonded length deformation 

(indicative of load transfer within the bonded zone). Figures 45 

and 46 show long-term relative displacement within the bonded zone 

1-17 
in a manner similar to that presented previously in Figures 19 

through 22. As before, the long rod-short rod difference and the 

wire 5-l difference are indicative of elongation (for positive 

differences) of the entire bonded zone. Good agreement was 

obtained between the long rod-short rod difference and the wire 5- 

1 difference for three of the four primary anchors (the exception 

being Anchor 51-upper). For the three anchors for which agreement 

was obtained, the elongation of the entire bonded zone was 

approximately 0.1 inch in each case. As before, elongation of the 

bonded zone is directly related to area under the load 

distribution curve. Thus, the additional long-term elongation 

indicates relatively less load transfer near the top of the bond 

zone and relatively increased load transfer deeper within the bond 

zone. 

6. DISCUSSION OF WALL PERFORMANCE 

The object of the discussion within this report section is to 

bring together the instrumentation results and use them to 

evaluate the wall design. 

Maximum outward movements of the wall have been approximately 3/4 

to 1 l/4 inches which corresponds to 0.2 to 0.3 percent of wall 

height, respectively. This range of movement appears to generally 

conform with that anticipated for the active soil pressure 

condition, for which the wall was designed. Soil movement has 

apparently occurred some distance behind the wall, creating an 

outward movement of the bonded zone of the anchors. 

In the long term (eight months), seven of the instrumented anchors 

have maintained their lock-off loads, two have increased somewhat 

(approximately 10 kips), and one has decreased somewhat 

(approximately 15 kips). Table IV shows the anchor load after 

eight months with respect to design loads. In that the anchor 

loads have not significantly increased above the design loads 

appears to indicate that the design loading conditions and 

resulting design tieback loads were not significantly 

underestimated. Had this been the case, the loading in the 

anchors would have increased to withstand these higher loads. 

This discussion on long-term tieback load with respect to design 

loading assumptions is generally true for the three anchors which 

experienced somewhat greater load variation. For the two anchors 

which experienced an increase in load (Anchors 66-lower and 67 

lower), it is possible that the design somewhat underestimated the 

actual loading. However, it is also possible that the design 

assumptions overestimated resistance of the pile embedment in this 

area, thus requiring the lower anchors to carry additional load. 

For the anchor which experienced a decrease in load (Anchor 67- 

upper 1 two possibilities exist. One is that the anchor bonded 

zone displaced an additional amount toward the wall in addition to 

1-18 
the general outward movement in the wall and anchor. A second is 

that the pile moved inward due either to stressing of the 

intermediate anchor, or possibly a redistribution of load toward 

the bottom of the wall (as reflected in increased load in Anchor 

67-lower). However, the second possibility is not supported by 

either Anchor 65-upper or 66-upper both of which had an 

intermediate' anchor below them. In addition, Anchor 66-lower 

experienced a load increase similar to that of Anchor 67-lower 

without a decrease in load in Anchor 66-lower. 

An intermediate level of tiebacks was required at the northern 

instrumentation section (piles 65, 66, and 67) because of failure 

of the upper level of anchors to achieved proposed design loads. 

The design loads for these intermediate anchors were calculated in 

the same manner as the original design loads, including the actual 

load for the upper level of anchors. Thus, while long-term load 

change data is not available for the intermediate anchors, it 

appears that the conclusions within this section wall 

performance and the relation to wall design are also appl?:able to 

this case. 

7. INSTRUMENTATION RECOMMENDATIONS 

The instrumentation on this project has generally provided the 

intended information. While some problems have surfaced, these 

are to be expected in an instrumentation program for which there 

was little precedent. The successful implementation of the 

program required intensive coordination between ourselves, the 

instrumentation consultant, the tieback contractor, the instrument 

suppliers, and the overseeing agencies. 

the instrumentation 

The customized design of 

necessitated many iterations between these 

parties. The need for this coordination cannot be overemphasized. 

The ability to select and procure the proper instrumentation on a 

"preferred sole source" basis was critical to the successes of the 

program. If similar instrumentation is planned on future projects 

we recommend that the 

services 

work be performed under a professional 

contract rather than be subjected to bid as part of the 

construction contract. 

The evaluation of instrument performance must be judged based on 

experience during instrument installation, proof testing of the 

tieback, -and long term behavior. The performance of 

instrument 

specific 

types are discussed below with recommendations for 

future consideration. 

Load Cells: The load cells have performed well and their 

reliability and durability have been excellent. The load cells 

had spherical bearings at their ends to increase accuracy. These 

bearings made difficult alignment of the cell, as well as 

alignment of the loading jack and chair. used for 

load testing, 

The equipment 

when mounted, cantilevered approximately four feet 

1-19 
beyond the retaining wall and therefore induced a moment at the 

spherical bearing location. Methods to stabilize the assembly 

during testing included very high alignment loads, supporting the 

jack with an overhead crane, and bolting the assembly to the 

soldier pile with long threaded rods and nuts. A convenient 

method of assembly was never developed. Future considerations to 

handle this problem could be: omission of the spherical bearings; 

or a "chair" similar to the one used for the jack which could be 

modified to support the load cell also. Bracing from the soldier 

pile and/or the ground below could be used for additional support. 

Connection of the wire leads to the load cell was inconvenient. 

On future projects each of the load cells should have a 

permanently installed lead wire with environmental connectors. 

Rod Telltales: Installation of the rod telltales was generally 

straight forward. For the short rod telltale on two of the 

intended secondary anchors it was impossible to insert the rod 

through the tube to the telltale anchor. The rod for one telltale 

could be inserted only a short distance, presumably because grout 

had entered the tube either at the telltale anchor connection, at 

a break in the tube, or at the top cap. In remaining 

installations this top cap was glued in place and cut off after 

grouting. The rod for the second unsuccessful telltale could be 

inserted almost to the telltale anchor. It was concluded that the 

tube had be come disconnected from the telltale anchor at their 

threaded connection. This connection could be improved in future 

installations by providing a longer threaded section. 

Twisting of the tendon and attached rod and wire telltales was 

recognized as a potential problem prior to initiation of the 

instrumen-tation program. This twist is believed to be a major 

cause of the problems when interpreting aspects of the wire and 

rod telltale data. After attachment of the telltales and 

placement of the tendon in the anchor hole, there was no way to 

maintain alignment of the instruments during grouting. As the 

grout was pumped, the casing was twisted and pulled. Difficulty 

in inserting the rod into the rod telltale sleeve is believed to 

be evidence of significant twist. However the actual degree of 

rotation experienced by the telltales and tendon is unknown. We 

believe there is no way to alleviate this problem in future 

installations involving stranded tendons. 

As discussed previously slippage may have occurred at the 

connection between the short rod telltale and the tendon. This 

possibility was recognized during the design phase, and 

substantial precautions were taken, including machining the base 

of the anchor to match the tendon spiral, gluing the two surfaces 

together, double-banding around the anchor and tendon, using a 

high tension banding system, and placing a soft material in front 

of the telltale anchors. Improvements should be sought in future 

installations, but space limitations may prevent a solution to the 

problem. As a minimum, the telltale anchors could be notched, 

l-20 
where the steel bands hold it to the strand, to improve mechanical 

interaction between them. Three bands, instead of two, would make 

the shear strength of the telltale anchor-strand connection 

tighter. Additional soft material placed in front of the anchor 

could help. alleviate the forces causing slippage. However, 

slippage may be impossible to prevent, particularly in the upper 

portion of the bonded zone where high loads in the tendon create 

the greatest possibility for differential movement between the 

tendon and grout. 

The telltale load cell (short rod telltale) has, in three of four 

cases, compared well with data from the load cell in providing 

long-term load change data, Thus, the telltale load cell appears 

to be a relatively reliable indicator of long term load change. 

In the one case where disparity exists it is not possible to 

determine which is correct, although it would be generally assumed 

that the load cell is more accurate. Where space within the 

anchor allows, two telltale load cells attached to different 

strands could provide redundancy in future installations. 

Wire Telltales: The wire telltales were used to measure 

deformation and thus estimate load transfer within the bond zone. 

The wire telltales gave more problems than the other 

instrumentation both in terms of installation, reading, and 

performance. In comparisons between rod and wire telltale data, 

good agreement was achieved in some cases while in other cases the 

wire telltales appears to have overestimated deflection. Several 

of the other wire telltales appear to have yielded unreasonable 

information. 

Small collars attached to the top of the wire telltales allowed 

the indicator to grip the telltale. The system was identical to a 

proven multi-point wire extensometer system which has performed 

well on other projects. However, when the range of the indicator 

was exceeded, it was necessary to move these collars. Because the 

wires were relatively stiff, they crimped easily during placement 

of the collars and made this operation difficult to perform 

properly. It is possible that some of the collars 

slipped during the 

may have 

load test and long term readings. In future 

installations, an improved method of collar attachment should be 

devised. Alternately, 

with the indicator, 

by use of spacers or shims in conjunction 

it might be possible to eliminate the need for 

collar relocation. 

The Peter Smith Mark I indicator used to read the wire telltales 

performed poorly. An increased reading range would help alleviate 

the problem discussed above. 

tensioning load be transmitted 

The indicator also required that the 

through the micrometer used to 

measure deflection. Thus, the tensioning load was significantly 

limited by the strength of the micrometer threads. An instrument 

with a larger range of 

reading accuracy, 

tensioning loads would provide greater 

wires 

by allowing a check on the functioning of the 

to be performed by a comparison of readings made at low and 

1-21 
high tensions. This multi-tension check was part of the original 

instrumentation proposal. The inability to perform these multitension checks reduced the reliability of the wire telltale data 

on this project. Such checks should be included as part of future 

wire telltale instrumentation. 

As discussed previously, slippage of the wire and rod telltale 

anchors may have also contributed to instrument error, and the 

measures recommended earlier for rod telltale anchors should also 

be adopted for wire telltale anchors. Where space within the 

anchor allows, we recommend additional telltales close to the top 

of the bonded zone. These would provide additional reference 

points should the question arise of slippage of the first telltale 

in the bonded zone. In addition, higher tensile loads exist in 

this zone than deeper in the anchor. This creates greater tendon 

elongations which are easier to detect amidst the normal data 

variation. 

An assumption common to both the wire and rod telltales is that 

deformations measured in one strand are representative of the 

entire tendon. General long term agreement in the telltale load 

cell (short rod telltale) with respect to the load cell tend to 

support the assumption. Instrumentation of duplicate strands, 

which would provide insight into this assumption, is typically 

limited because of space limitation. 

The success of the use of wire telltales on a stranded tendon to 

estimate load transfer within the bonded zone on this project is 

in question because of possible anchor slippage. In addition, 

certain inaccuracies exist in the remainder of the data. The 

problem of tendon twist during the pulling of casing may be the 

cause of these inaccuracies. In that case, the problem of twist 

may be impossible to overcome. However, the implementation of the 

improvements suggested within this section should improve the data 

quality and could yield more successful results. 

Inclinometers: The inclinometers appear to have provided 

consistent and reliable information. Although examination of the 

data indicates that base fixity was not quite obtained, the 

deviations were small and do not appear to have caused significant 

error in deflection measurements. 

Survey: Survey appears to have provided the reliable data 

anticipated from this proven instrumentation technique. The 

survey points were installed, read and reported by the Georgia DOT 

crews. Thus, credit for success should belong to them. As should 

be expected, many points on the top of the wall were destroyed 

before significant data could be obtained. 

Lonq-Term Readinas: We recommend continued long-term readings of 

load cells, short and long rod telltales, inclinometers, and 

survey points. Long-term reading of wire telltales #l and #5 

should also be considered to provide redundancy for the rod 

l-22 
telltales. If this is performed, readings for wire telltales #2, 

#3, and #4 could also be continued and might provide information 

concerning long-term redistribution of load transfer within the 

,bonded zone. 

8. CONCLUSIONS 

The following are our primary conclusions concerning the permanent 

tieback retaining wall, Georgia DOT Project P.E. l-75-02 (41). 

Instrumentation Results 

Conclusions which may be drawn from the test data are summarized 

in the following points: 

1. During the load test, the bottoms of 

instrumented anchors showed little or no movement. 

the primary 

Estimation of 

modulus values for the anchor free length may have been affected 

by slippage of the short rod and wire #l telltale anchors. 

2. The question of telltale anchor slippage during the load 

test also affects the estimation of load transfer within the 

bonded zone. If slippage did not occur, 

transfer appears to have occurred in 

relatively high load 

bonded 

the upper portion of the 

zone. If slippage did occur, estimated load transfer 

appears more reasonable. 

3. Comparison of jack pressure gage loads with load cell 

results indicates that jack gage overestimates load during loading 

and underestimates load during unloading. Load as measured by the 

jack has an apparent accuracy of + 5 to 10 percent. 

4. Significant variation in load at the anchor head has not 

occurred for seven of the ten instrumented anchors. Two of the 

other anchors have gained load (approximately 10 kips) while one 

has lost load (approximately 15 kips). 

5. Most of the measured wall deflection occurred within one 

to two months after excavation. 

that deflection has stabilized. 

Long-term measurements indicate 

The magnitude of both horizontal 

and vertical deflections appear to be related to wall height, and 

are well within the range of past experience. 

6. The anchors' 

approximately 

bonded zones were displaced a distance of 

l/2 to 3/4 inch during outward deflection of the 

wall and anchor head. result of overall 

soil deformation. 

This appears to be the 

Some load redistribution within the bonded zone 

occurred during this displacement. 

l-23 
Wall Performance 

Based on the results of the load testing and long-term load and 

displacement performance monitoring to date, it appears that the 

permanent tieback wall is performing satisfactorily. The 

assumptions and procedures used to design the wall and the 

construction methods used to accomplish the design also appear 

satisfactory. Thus, it can be stated that the permanent tieback 

retaining wall system is a suitable design and construction 

alternative for future highway construction in similar soils in 

the State of Georgia and in the southeastern United States. 

It is recognized that one of the major concerns of owner-agencies 

is the anchor permanency against long-term corrosion and/or creep; 

or: "What is the design life of the anchor wall?" This concern 

can be addressed by relating the past 25-40 years of experience in 

the use of permanent tieback walls to the design, installation and 

testing procedures used in this project. 

The design life of anchor permanency can be ensured through 

establishing the technical feasibility for a permanent tieback 

wall for the specific site, an evaluation of the risk by the 

owners based on the feasibility study, the pre-qualification of 

designers and contractors, the selection of tieback type including 

state-of-the-art corrosion protection, and the establishment of 

load testing acceptance criteria and long-term monitoring. 

To estimate the design life of this particular permanent anchor 

wall, one must evaluate the short-term load holding capacity of 

the anchors, the long-term load holding behavior which is eight 

months to date, and consider the design and construction 

procedures. The test results presented in this report indicate 

that the wall is behaving as designed. The design and 

construction was carried out by qualified professionals, and the 

corrosion protection system selected and installed is assumed to 

be providing the long-term corrosion performance for which it was 

designed. Therefore, the life expectancy of this permanent 

tieback wall can be estimated to be comparable to a reinforced 

concrete structure designed and constructed with similar 

professional care. 

Safe Desian Loads 

Ultimate capacity of the instrumented anchors is shown in Table 

IV. Ultimate loads obtained on this project typically ranged from 

four to five kips/foot of bond length. Anchor loads for the 

subject wall were reduced from these ultimate loads by use of a 

safety factor. The data for the instrumented anchors and overall 

wall system indicate that under similar subsurface and 

construction conditions, preliminary safe design loads on other 

projects could be derived from these ultimate loads and a similar 

safety factor. Of course, actual loads must be confirmed with 

load testing. 

l-24 
INSTRUMENT TYPE 

Inclinometer 

Load Cell 

Wire Telltales 

TABLE I 

SUMMARY OF INSTRUMENTATION 

MANUFACTURER 

Slope Indicator, Co. 

Brewer Engineering 

Laboratories, Inc. 

Geokon, Inc. 

PARAMETER 

MEASURED 

Horizontal deformations of wall face 

and retained soil 

Load in anchor 

Load distribution 

in anchor bond zone 

EQUIPMENT (MAJOR PARTS) 

. Inclinometer system with 

Digitilt sensor, LCD 

indicator, 2.75" o.d. 

plastic casing installed 

using grout backfill. 

. 225-kip capacity with BLR 

120 Indicator Box, and 

digital readout, bearing-- 

washers and spherical 

bearing plates. Gage and 

calibrated pull bar to 

re-establish zero reading 

at any time. 

s/s wire type 16 G.A. 

1 l/4" oil-filled nylon 

tube around each wire 

. 5/8" grooved s/s anchor 

attached to the strand 

using clamps and epoxy 

. 3/4" pvc pipe on the axis 

of the entire anchor with 

slots for grout entry and 

nylon tube entry. 

. Peter Smith Mark I extensometer readout. % 

. Readings made using a ' 'd 

chair attached to the 3 anchor head. l-l x 

P 
TARLE I (can't.) 

SUMMARY OF INSTRUMENTATION 

INSTRUMENT TYPE MANUFACTURER PARAMETER EQUIPMENT (MAJOR PARTS) 

MEASURED 

Rod Telltales Geokon, Inc. Load chanqe in . 

anchor 

Provides data 

redundancy in bond . 

zone 

. 

. 

r 

+ 

10 Survey 

. 

Flush-coupled l/4" s/s 

rod within an oilfilled PVC pipe. 

Anchors same as wire 

telltale 

l-7/8" s/s reference 

plate required on the 

anchor head. 

s/s chair for in-situ 

calibration 

micrometer for readout 

Provided by Georgia Horizontal and . Electronic distance 

DOT vertical defor- meter (EDM) for horizonmation of wall zontal deflections 

face and top of . conventional leveling 

retained soil techniques. 
TABLE II 

SUMMARY OF ACCEPTANCE CRITERIA 

CREEP TESTING 

INSTRUMENTED ANCHORS 

ANCHOR 

Creep/Cycle LENGTH OF TEST 

(inches) (minutes) 

51 upper 0.003 

51 lower 0.020 

66 upper 0.001 

66 lower 0.002 

67 upper 0.008 

67 lower 0.017 

65 upper 0.029 

65 lower 0.009 

52 upper 0.013 

50 upper 0.008 

The creep per log cycle computed by: 

d2 - dl 

creep/cycle = 

5”: 

5 

50 

5 

10 

lo" 

5 

5 

d - deflection 

T- time 

1-A-3 
TABLE III 

APPARENT MODULUS OF UNBONDED TENDON LENGTH 

DETERMINED DURING LOAD TESTS 

Short Rod Telltale Data Wire Telltale #I Data 

E Loading E Unloading E Loading E Unloading 

PRIMARY 

ANCHORS 

(Psi) (Psi) (psi> (Psi) 

51-lower 21.4~106 30.7~106 21.9x106 29.2x106 

51-upper 22.4~106 26.3~106 22.4~106 29.5x106 

66-lower 22.7~106 28.9x106 26.3~106 27.3~106 

66-upper 19.3x106 30.6~106 20.0~106 29.8x106 

SECONDARY 

ANCHORS 

67-lower 23.9x106 25.7~106 

67-upper 24.0~106 23.9x106 

52-upper 24.2~106 25.2~106 

50-upper 19.3x106 36.2~106 

65-lower 22.4~106 32.8~106 

65upper 26.2~106 30.49106 

l-A-4 
TABLE IV 

SUMMARY OF INSTRUMENTED ANCHOR PERFORMANCE 

LOAD AFTER LOAD AFTER CREEP 

BOND ULTIMATE LOAD DESIGN SETTING EIGHT KIPS PER 

LENGTH ~XIPSPER LOAD WEDGES MONTHS LOG 

ANCHOR (FEET) (KIPS) FOOT (KIPS) (KIPS) (KIPS) CYCLE 

51-u 

51-L 

66-U 

66-L 

50-u 

52-u 

65-U 

65-L 

67-U 

67-L 

32 189.9* 

35 

32 

35 

33 

33 

32 

35 147.2 

32 

35 

188.0 

131.9* 

155.0 

190.5 

201.7 

114.2* 

130.2* 

158.7 

5.9 132.6** 128.4 132.9 tt 

5.4 125.3 126.0 126.0 3.9 

4.1 91.8** 94.1 92.3 tt 

4.4 104.1 99.9 108.9 0.5 

5.8 129.7 147.6t 145.6 1.9 

6.1 138.8 154.6.t 152.4 3.4 

3.6 78.4** 83.0 79.3 tt 

4.2 98.0 94.7 94.1 0 

4.1 91.1** 94.2 79.2 0.7 

4.5 105.8 106.1 117.3 ttt 

NOTES: 

Grout pressure of approximately 70 psi used on all anchors 

* Approached acceptance criteria before reaching predetermined 

ultimate load; other anchors achieved predetermined load 

without approaching criteria. 

** New design load based on results of load test. 

t Lock-off Load (load after setting wedges not determined) 

tt bad variation judged too erratic for accurate creep determination. 

ttt Anchor gained load. 

l-A-5 
-N 

ATLANTA 

COCA COLA 

1 

BOTTLING 

COMPANY 

l-75 CONSTRUCTION -I 

LIMITS TIEBACK WALL 

PLAN ‘INSTRUMENTATION 

STATIONS 

CANTILEVER WALL PERMANENT TIEBACK WALL 

ELE” 

IN FT. TOP OF WALL 9 

960 

920-I L BOTTOM OF WALL 

(LAGGING & FACING) 

PROFILE 

FIGURE 1 -PLAN AND PROFILE OF PERPAANENT TIEBACK WALL 

l-A-6 
8 

51 

-N 

LEGEND CURB LINE 

. PERMANENT TIEBACK ANCHOR 

. PRIMARY INSTRUMENTED ANCHOR (PA) 

,$ SECONDARY INSTRUMENTED ANCHOR ,SA) 

7 a SLOPE INCLINOMETER (SI) 

WILLIAMS STREET / 

SI I 

8 ! 

7 

I 

: 

SOLDIER PILE CURB LINE 

b 

-; 

SI ? 

I 

II II II 

4 II - 

PLAN 

(TIMBER LAGGING NOT SHOWN) 

::jj SOLDIER 

-L; PILE _/ 

3 CONCRETE 

FACING 

ELEVATION NOTE: LOWER SECONDARY ANCHOR 

(TIMBER LAGGING NOT SHOWN) 

INSTRUMENTATION (PILES 50 

& 52 DESTROYED) 

FIGURE 2 - PEFWANENT TIEBACK WALL INSTRUMENTATION 

l-A-7 
960 

950 

IW 

i.z 

z 940 

z 

0 

5 > 930 

W 

J 

W 

920 

910 

. PERMANENT TIEBACK ANCHOR 

3 

/ 

8 PRIMARY INSTRUMENTED ANCHOR (PA) 

“i 9 SECONDARY INSTRUMENTED ANCHOR (SA) 

n SLOPE INCLINOMETER (9) 

z 

1 

SOLDIER PILE 

WILLIAMS STREET 

SI 

l i 

CURB LINE 

( 

CONCRETE I 

FACING / I 

PLAN 

(TIMBER LAGGING NOT SHOWN) 

-- 

-- / SOLDIER PILE 

ELEVATION 

(TIMBER LAGGING NOT G-I~wN) 

L CONCRETE FACING 

FIC;URE 3 - PERMANENT TIEBACK WALL INSTRUfviENTATION 

l-A-8 
INCLINOMETER 

CASING 

(SI) -T-- 

LONG ROD 

TELLTALE 

WIRE 

TELLTALE 

SOLDIER PILE 

>I(/ JyQ 

SHORT ROD / 

TELLTALE A 

I 

(RT) u ‘GROUT/ 

LEGEND 

0 1 INSTRUMENT POSITION DESIGNATION L WIRE\ TFI I l-AI F . ---. , .-- 

(WT) 

NOTE: 

ANCHOR IS SHOWN IN FINAL LOCK-OFF 

CONDITION. 

FIGURE 4 - SCHEMATIC OF PRIMARY INSTRUMENTED TIEBACK ANCHOR 

l-A-9 
WIRE TELLTALE 

) )-PVC PIPE 

i?nn TFI I TAI F 

CORRUGATED 

PLASTIC PIPE 

FlGUR: 5 - CROSS SECTION OF INSTRUMENTED TIEBACK ANCHOR 

1-A-10 
I ,...“R --. 
50 MlNUTti 

CREEP TEST--\ TEST-- 

WIRE TELLTALE 

145.6 KIPS 

ACCEPTANCE 

CRITERIA 

WIRE TELLTALE #5 

NOTE: 

AFTER SETTING WEDGES 

LOAD= 126.0 KIPS 

FIGURE 7;- LOAD DlSPLACElZIiENT CURVE LOAD TEST 

PILE 51. LOWER ANCHOR 

- -I_- 

FREE LENGTH (LOAD TEST) = 19.6 FT. 

BOND LENGTH = 34.6 FT. 

---- 

-- 

I----- __ ,----. ..--,------, 

I.6 I .n 2.0 2.2 
150 

14; 

123 

ii 100 

Y 

: 

2 80 

50 

40 

20 

/ 5 MINUTE 

/ 

// WIRE TELLYLE #I\ 

CREEP TEST 

SHORT ROD TELLTALE 

- LVI.V VT”” 

/ 

DIAL GAGE 

/’ 

WIRE TELLTALE #5 

‘TANCE CRITERIA 

TELLTALE 

SEATING LOAD = 31.6 KIPS 

/ 

/ PILE 66. UPPER ANCHOR 

FREE LENGTH (LOAD TEST) = 22.6 FT 

BOND LENGTH = 31.7 FT. 

102.5 KIPS 

NOTE: 

AFTER SETTING WEDGES 

LOAD= 94.1 KIPS 

0 9.2 0.4 0.6 0.9 1 .o 1.2 1.4 I .6 1.6 

DEFLECTION, INCHES 

FIGURE 8 LOAD DISPLACEMENT CURVE - LOAD TEST 
50 MINUTE 

CREEP TEST CREEP TEST 

120 SHORT ROD TELLTALE 

100 

In WIRE TELLTALE X5 

ii 

j2 ACCEPTANCE 

Ii 00 

CRlTERlA -;y/ /s$fi Ly R&;LLTAL, 

. 

/ 

60. /' 

/ 

-' 

NOTES: 

AFTER SETTING WEDGES 

LOAD=99.9 KIPS (a STRANDS 

ONLY! 

UNLOADING AND RELOADING 

FOR WIRES #l AND #5 ARE 

NOT INCLUDED 

PILE 66, LOWER ANCHOR 7. , 

k 

FREE LENGTH (LOAD TEST) = 19.1 FT. 

SEATING LOAD = 10.0 KIPS BOND LENGTH = 34.4 FT. 

0 I 1 --- 

, I I I 1 1 

0 0.2 0.4 0.6 0.6 I.0 1.2 I .b I .6 1.6 2.0 

DEFLECTION, INCHES 

KIPS 

FIGURE 9 LOAD DISPLACEMENT CURVE - LOAD TEST 
\ 

\ a r 

a 111 

a 

Li 

‘3 \ 

a 

. 

N 

W 

. . 

(u 

d 

. 

N 

N 

. 

N 

0 

. . 

N 

co 

. . 

ln 

wow 

‘, I 

u 

Z 

*Z 

‘, 0 

iz 

ii 

J 

NLL ‘, w 

cl 

0 

_ 

m 

. 

0 

W 

_ 

0 

* 

0 

N 

_ 

0 

-0 

SdlN 6 avo7 

1-A-15 
\ i \ 0 
N 

0 

c3 

d 

5 

W 

6 

P 

d 

N 

d 

0 

.O 

I I I I 

z iti 

0 0 0 W P .O 

N m 

SdlM * avow 

1-A-17 

N 

W 

-I 
\ 

a 

\h \ a 
r * 

a \ 

W L 

r W 

a 

0 0 

P 0 

N 0 

ID q 

Sdl>l ' avoi 

1-A-19 
200 

180 

160 

140 

120 

PO 

60 

40 

20 

0 

10 MINUTE 

CREEP TEST- 1 

SHORT ROD TELLTALE 

/ 

-SEATING LOAD = 5.0 KIPS 

NOTE: UNLOADING AND RELOADING ACCORDING TO 

DIAL GAGE ARE NOT INCLUDED FOR CLARITY. 4, / 

I I I I I I I I I I I I I I 

0 G.2 0.4 0.6 0.8 1.0 1.2 1.4 1.6 1.8 2.0 2.2 2.4 2.6 2.8 

DEFLECTION , INCHES 

FIGURE 15-LOAD DISPLACEMENT CURVE - LOAD TEST 

PILE 67 , LOWER ANCHOR 

FREE LENGTH (LOAD TEST) = 19.7 

BOND LENGTH=JS.O FT. 

/ 

:RI 

NOTE: 

AFTER SETTING WEDGES LOAD 106.1 KIPS 

A 

FT. 
LOG TIME (MINUTES) 

; 40 

g 30 

w 

0 20 

f 10 

1 

0 

0 

0 o/-----o 

/- -- 

000 O O 

I I I 

50 

I 

100 

1 

150 200 250 

LOAD , KIPS 

LEGEND: 

20 ANCHOR NUMBER 

(13oh) LOAD, KIPS 

FOR FURTHER INFORMATION 

ON THE FORM AND INTERPRETATION OF THE DATA PRESENTED IN THIS FIGURE, 

PLEASE REFER TO REPORT 

NO. FHWA!RD-61/150, 

PAGES 91-95 

FIGURE lEj- CREEP CURVES FROM PERFORMANCE TESTS 

1-A-21 
19c 

160 

16C 

14c 

120 

60 

60. 

40, 

20. 

oPILE 51, UPPER ANCHOR 

WIRE 5 - I 

LONG ROD 4 

-0.1 74--- 0 +0.05 +0.1 

RELATIVE DEFORMATION, INCHES 

1 - 

ROD 

PILE 51, LC 

WIRE 1 - 

SHORT ROD\ 

1 I I 1 

-0.2 -0.15 -0.1 -0.05 

IER ANCHOR 

\ 

I 

I 

-WIRE 5 - \ LONG ROD 

\ 

I 

\ 

LONG ROD - 

DIAL GAGE I 

I 

4 

I 

I 1 , 

+0.05 +0.1 +0.15 *0.2 

RELATIVE DEFORMATION. INCHES 

I 

-t 

-4 

-2 

190 

160 

160 

140 

120 

2 

2 IO0 _ 

2 

s 

30 

$0 

0 

0 

I 0 

FIGURE 17- COMPARISON OF WIRE TELLTALE VERSUS ROD TELLTALE 

DURING LOAD TEST 

1-A-22 
163 

140 

120 

100 

60 

40 

20 

0 

PILE 66, UPPER ANCHOR 

1 - 

ROD 

I 

-0.1 -0.05 !- 0 +0.05 +0.1 

RELATIVE DEFORMATION, INCHES 

PILE 66, LOWER ANCHOR 

t . 

\ . 

t . 

1 

i 

t I 

I 

. 

t LONG ROD- . 

DIAL GAGE 

1 

RE 1 - 

RT ROD 

WIRE 5 - 

LONG ROD 

-0.1 ----I-- -0.05 0 +0.05 +0.1 

RELATIVE DEFORMATION, INCHES 

L 

.I60 

‘140 

,120 

60 

40 

20 

0 

FIGURE IS-COMPARISON OF WIRE TELLTALE VERSUS ROD TELLTALE 

DURING LOAD TEST 

1-A-23 
190 

180 1 

-1 160 

L-WIRE 2-1 

/ WIRE 3-l 

7 

L LONG ROD - SHORT ROD 

LOAD 

WIRE 5-f 

0 I 1 I I I I 

0 0.1 

1 

0.2 u.3 0.4 0.5 0.6 0.7 

RELATIVE DEFORMATION . INCHES 

FIGURE 19 - RELATIVE DEFORMATION IN BOND ZONE - LOAD TEST 

l-A-24 

PILE 51 - UPPER ANCHOR 

LOAD TEST 
180 

160 

140 

120 

8C 

60 

4 

/ i I 

i 

\ 

i 

t 

:L- WIRE 2-1 

i 

i 

i / ’ k- WIRE 4-1 

LONG ROD - 

SHORT ROD7 

i 

/ 

/ 

I I I 

i 

I 

: 

4 

I 

I 

: 

- 11.4 KIP SEATING LOAD 

WIRE 5-l 

P 

PILE 51 - LOWER ANCHOR 

LOAD TEST 

0 0.1 0.2 0.3 0.4 0.5 0.6 

RELATIVE DEFORMATION , INCHES 

FIGURE 20 - RELATIVE DEFORMATION IN BOND ZONE - LOAD TEST 

l-A-25 
‘L LONG ROD - SHORT ROD 

NOTE: 

WIRE COMBlNATlON 2-l MEASUREMENTS 

GIVE NEGATIVE DEFLECTIONS. 

- NEGATIVE VALUES IN THIS 

RANGE NOT SHOWN 

t 3t.6 KIP SEATING LOAD 

PILE 66 - UPPER ANCHOR 

LOAD TEST 

4-T---- c.20.3 0.4 0.5 

RELATIVE DEFORMATION , INCHES 

FIGURE 21 . RELATIVE DEFORMATION IN BOND ZONE - LOAD TEST 

1-A-26 
16C 

140 

120 

1 oc 

kr! 

2 

Ii 

80 

6C 

ItI3- 

0-l 

1 

D ! 

c 

I I I I I 

\ 0.1 012 0.3 04 0.5 0.6 

RELATIVE DEFORMATION INCHES 

i WIRE 3-f I 

-WIRE 4-1 

FIGURE 22- RELATIVE DEFORMATION IN BOND ZONE - LOAD TEST 

1-A-27 
BOTTOM 

,,,‘,“,“,‘,-/- 7.5 FT. ;I, 7.5 FT. _ 1 _ 

OF ANCHOR 

7.5 FT. 

NOTE: 

LOAD !S IN ADDITION TO SEATING LOAD. LOAD IN FREE 

LENGTH ASSUMED = 150 KIPS [ABOVE SEATING LOAD) 

TOTAL CALCULATED ELONGATION 

OF BONDED ZONE=O.BO IN. 

TOTAL CALCULATED ELONGATION 

OF BONDED ZONE = 0.60 IN. 

LEGEND 

@ -INSTRUMENT POSITION DESIGNATION (SEE FIGURE n) 

AL -CALCULATED ELONGATION OF SEGMENT OF BONDED 

ZONE (INCHES) 

BOTTOM 

OF ANCHOR 

TOTAL CALCULATED ELONGATION 

OF BONDED ZONE = 0.90 IN. 

TOTAL CALCULATED ELONGATION 

TOTAL CALCULATED ELONGATION 

OF BONDED ZONE = 0.90 IN. 

OF BONDED ZONE = 0.60 IN. 

A. HYPOTHETICAL LOAD TRANSFER DIAGRAMS WITH CALCULATED 

TOTAL BONDED ZONE ELONGATION = 0.60 INCH 

6. HYPOTHETICAL LOAD TRANSFER DIAGRAMS WITH CALCULATED 

TOTAL BONDED ZONE ELONGATION = 0.90 INCH 

FIGURE 23 - HYPOTHETICAL LOAD TRANSFER DIAGRAMS AND CALCULATED TOTAL ELONGATION OF BONDED ZONE 
FREE 7.5 FT. BOTTOM OF 

_ 7.5 FT. ANCHOR 

LENGTH 

ANCHOR 66-UPPER 

LEGEND: 

0 -.- LOAD CELi=79.4 KIPS 

ISO1 oo- 

:: 

z 

li 

ti 

soO( 

LOAD CELL= 109.4 KIPS 

LOAD CELL-131.9 KIPS 

ANCHOR 66-LOWER 

LEGEND: 

0 -.- LOAD CELL179.7 KIPS 

A------ LOAD CELL-1 19.9 KIPS 

a-..-- LOAD CELL=139.3 KIPS 

NOTE: POINTS REPRSENT CALCULATED AVERAGE LOAD WITHIN 7.5 FT. SEGMENTS 

BASED ON RELATIVE DEFLECTIONS AS MEASURED BY WIRE TELLTALES 

AND ADJUSTED FOR ASSUMED SLIPPAGE OF WIRE +l 

CURVES REPRESENT LOAD DISTRIBUTIONS FOR WHICH CALCULATED BONDED 

ZONE ELONGATION IS EQUAL TO THE MEASURED ELONGATION ADJUSTED 

FOR ASSUMED SLIPPAGE OF WIRE trl 

‘DlsTRlBuTl0~ AT SEATING LOAD is ASSUMED 

FIGURE 24 - LOAD DISTRIBUTION IN BONDED ZONE (LOAD TEST) ~SsuitiiNG SLIPPAGE 

OF WIRE * 1 AND SHORT ROD TELLTALE ANCHORS 

1-A-23 
115- 

110- 

d! 

0 

0 

‘i 105 - 

si 

OW 

JV 

yu2 

a0 0 0 

-3J 100 

PILE 50 

1 SOtJACK 

LOAD TEST 

95- 

90 I I I 1 

0 50 100 150 200 

LOAD CELL, KIPS 

LEGEND: 

-0 LOADING 

3 

UPPER 

-.A UNLOADING 

FIGURE 25 

~OWl!‘ARISON OF LOAD CELL AND HYDRAULIC JACK 

DURING LOAD TEST 

1-A-30 
115 

110 

105 

100 

95 

90 

PILE 51 

300tJACK 

LOAD TEST 

0 50 100 150 200 

LOAD CELL, KIPS 

LEGEND: 

-0 LOADING 

3 

UPPER 

-)4 UNLOADING 

4 L3ADING LOWER 

-.4 UNLOADING 

FIGi;RE 26 

SO!RPARISOIJ OF LOAD CELL AND HYDRAULIC JACK 

DURING LOAD TETT 

1-h-31 
-*- ,-.y-- .-’ 

PILE 52 

1 SOtJACK 

LOAD TEST 

90 I I I 

0 

1 

50 100 150 200 

LOAD CELL. KIPS 

LEGEND: 

___O LOADING 

3 

UPPER 

--4 UNLOADING 

FIG’,rRE 27 

COQl!‘ARISON OF LOAD CELL AND HYDRAULIC JACK 

DURING LOAD TEST 

1-A-32 
11: 

11c 

105 

100 

95 

90 

PILE 65 

1 SOtJACK 

LOAD TEST 

I I I 

50 100 

I 

150 200 

LOAD CELL 8 KIPS 

LEGEND: 

0 LOADING 

-. 4 UNLOADING 3 

UPPER 

a LOADING 

-.A 3 

LOWER 

UNLOADING 

FIGURE 28 

COV’/IPARlSON OF LOAD CELL AND HYDRAULIC JACK 

DURING LOAD TEST 

1-A-33 
115 

110 

8 

j; 

,“i 105 

OW 

;:: 

!i; 

100 

95 

90 

PILE 66 

\ 300tJACK 

+ 

LOAD TEST 

0 50 100 150 

LOAD CELL. KIPS 

200 

LEGEND: 

,-O LOADING 

f 

UPPER 

--.d UNLOADING 

d LOADING 

3 

LOWER 

-.- UNLOADING 

FIGURE 29 

S:OMPA.RISON OF LOAD CELL AND HYDRACLIC JA3K 

CURING ‘-OAD TECT 

l-A-34 
105 

0 0 

PILE 67 

1 SOtJACK 

LOAD TEST 

LOAD CELL, KIPS 

LEGEND: 

4 LOADING 

UPPER 

-.a UNLOADING 

-I LOADING LOWER 

FlGC!RE 30 

COMDARISON OF LOAD CELL AND HYDRAULIC JACK 

DZRiNG LOAD TEST 

1-A-35 
PILE 51 , UPPER ANCHOR 

LOWER 

ANCHOR 

STRESSED 

1 

LOAD CEL>. 4 *. 

I I lllll //‘I I-----l 100 300 1000 

DAYS 

\ 

NOTE: AFTER SETTING WEDGES , 

/ i 

\ 

-5 LOAD= 128.4 KIPS 

J 

\ \ 

L 

PILE 51 , LOWER ANCHOR 

SHORT ROD TELLTALE 

NOTE: AFTER SETTING WEDGES , 

LOAD=1 26.1 KIPS 

FIGURE 31 . LONG TERM LOAD VARIATION 

l-A-36 
INTERMEDIATE 

ANCHOR 

STRESSED 

I LOWER 

+10 1 1 

ANCHOR 

PILE 66 . UPPER ANCHOR STRESSED 

iWIRE NO.1 

TIME , DAYS 

NOTE: AFTER SETTING WEDGES , 

LOAD=94.1 KIPS 

PILE 66 , LOWER ANCHOR 

WIRE NO.1 

LOAD CELL 

SHORT ROD 

TELLTALE 

TIME , DAYS 

I I I IllI I I I I IIll I 

I “1”‘l 

10 1OG 300 1 coo 

TELLTALE LOAD CELL CHANGES COMPUTED 

WITH AREA OF FOUR STRANlk ONLY 

NOTE: AFTER SETTING WEDGES. 

LOAD=99.9 KIPS 

FIGURE 32- LONG TERM LOAD VARIATION 

1-A-37 
LOWER 

ANCHOR 

STRESSED 

+5 - 

12 

E 

TIME , DAYS 

s 

2 0 

6 

2 

SHORT ROD TELLTALE 

s NOTE: LOCK-OFF LOAD=1 47.6 KlPS 

-5 - 

PILE 50 . UPPER ANCHOR 

LOWER 

ANCHOR 

STRESSED 

+5 1 

TIME , DAYS 

5 10 100 300 1000 

f O 

I I IIIIII I I ,,,,, I I 11,111 

6 SHORT ROD TELLTALE 

1 NOTE: LOCK-OFF LOAD=154.6 KIPS 

-5 

PILE 52 , UPPER ANCHOR 

FlGURE 33-LONG TERM LOAD ‘JARIATION 

1-A-38 
INTERMECIATE 

ANCHOR 

STRESSED 

1 , 

+5 

ii! 

E 1 TIME ( DAYS 

‘W 

k2 10 100 300 1000 

% n I I I Illlll I I I Illll I I lllll 

SHORT ROD TELLTALE 

t - 

2 

s 

LOWER 

ANCHOR 

STRESSED 

NOTE: AFTER SETTING WEDGES , 

LOAD=83.0 KIPS 

-5 -1 PILE 65 , UFFER ANCHOR 

TI:J;E , DAYS 

10 100 300 1000 

IQ 

o-w-1 ’ “1’ I 

v 

2 SHORT ROD TELLTALE 

s 

NOTE: AFTER SETTING WEDGES , 

LOAD=94.7 KIPS 

PILE 65 , LOWER ANCHOR 

FIG;;RE 34-LONG TERM LOAD VARIATION 

l-A-39 
TIME , DAYS 

NOTE: AFTER SETTING WEDGES , 

fc 

LOAD194.2 KIPS 

z -5 -’ 

INTERMEDIATE 

w 

is 

STRESSED 

9 

u 

ii 

-10 - 

A 

SHORT ROD TELLTALE 

-IS- 

+15 

0 

PILE 67 , UPPER ANCHOR 

SHORT ROD TELLTALE 

NOTE: AFTER SETTING WEDGES , 

LOAD=l06.1 KIPS 

I I 1 Illll I I I Illll I 

1 I ’ ““V 

10 100 3co 1000 

TIME . DAYS 

PILE 67 1 LOWER ANCHOR 

FIGURE 35-LONG TERM LOAD VARIATION 

1-A-40 
TIEBACK ANCHORS - 

‘,;/ 

_/_/_/)i__ 

/ 

/ 

Y’ ./ 

/ 

V’ 

- BOTTOM OF 

SOLDIER PILE 

DATE @F MEASUREMENT 

0 l-7-63 

@ i-26-62 

@ 3.2A-93 

@ 6-14-83 

0 6-26-63 

10 FEET 

HORIZONTAL SCALE 

? 5.5 1 INCH 

1 1 

DEFLECTION SCALE 

FIGURE 36 - HORIZONTAL DEFLECTION BEHIND PERMANENT 

TIEBACK WALL (PILE 51) 

l-A-41 
970 - 

960 -’ 

920 - 

/ 

I 

910--1 

51-66-2 

PATE OF MEASURWENT~ 

,D l-i-E3 

2 i-26-83 

z 3-20-63 

s 6-14-63 

g 0.Z&P3 

I- 

?&. 

‘7 .5 

i 

” 

! /I HORIZONTAL 

! iJ 

DEFLECTION 

-BOTTOM OF 

SOLDIER PILE 

-66-I 

9 5 10 FEET 

- 

HORIZ0NTP.L SCALE 

3 3.5 1 INCH 

I I I 

DEFLECTION SCALE 

FIGURE 37 - HORIZONTAL DEFLECTION BEHIND PERMANENT 

TIEBACK WALL (PILE 66) 

l-A-4 2 
0.25 I I I I I I I I I I 

g 

z -0.75 

~TH= 50 

.--v_ 

n, DEPTH = 22’ 

STRESSED -- --+J------y ?- UPPER ANCHOR 

W LOWER ANCHOR STRESSED 

5 

1992 1993 

-1.00 A- I I I I I I I I I 

E SEP OCT NOV DEC JAN FEB MAR APR MAY JUN JUL AUG 

9 

0 

z 

W 

G: 0.25 

W 

n 

2 0.00 

IR E -0.25 

-0.50 

-0.75 

-1 .oo 

INSTRUMENT NO. SI - 51 - 1 

I I I I I I I I I I 

I 

INSTRUMENT NO. SI - 51 - 2 

--a--- 

-\ DEPTH = 30’ 

-\ -+------Lj 

‘JPPER ANCHOR STRESSED 

DEPTH = 13’ 

a 

LOWER ANCHOR STRESSED 

1992 1993 

SEP OCT NOV DEC JAN FEE MAR APR MAY JUN JUL AUG 

NOTE: (+)MovEMENT AWAY FROM EXCAVATION 

(-)MOVEMENT TOWARD ExCAVATION 

FIGURE 38 - INCLINOMETER DEFLECTION VERSUS TIME 

l-A-43 
0.25. 

0.00 

-0.25 

G -0.50 

D 

ii 

-0.75 

!l! 

2 -1 .oo - 

SEP OCT NOV DEC JAN FEB MAR APR MAY JUN JUL AUG 

0 I I I I I I I I I 

INSTRUMENT NO. SI - 51 - 3 

LOWER ANCHOR STRESSED 

UPPER ANCHOR STRESSED 

z 

!i 

i: 

u 

W 

-I 

:: n 0.25 

;: 

g 0.00 

No 

g -0.21 

I 

-0.50 

-0.75 

-I .oo 

I I I I I 1 I I I I 

INSTRUMENT NO. SI - 66 - 1 

---_ 

DEPTH = 7’ 

UPPER ANCHOR INSTALLED 

LOWER ANCHOR INSTALLED 

SEP ocr NOV DEC JAN FEE MAR APR MAY JUN JUL AUG 

NOTE: (+)MOVEMENT AWAY FROM EXCAVATION 

klh40vEh1E~T TOWARD EXCAVATION 

FIGURE 39 - INCLINOMETER DEFLECTION VERSUS TIME 

l-A-44 
0.2; 

0.00 

-0.25 

-0.50 

-0.75 

-1 .oo 

T 

I I I I ’ ’ ’ ’ ’ ’ ’ 

INSTRUMENT NO. SI - 66 - 2 

DEPTH = 20’ 

---cr--L,- 

-‘-& --B-B-- 

DEPTH = 11’ 

UPPER ANCHOR 

INSTALLED LOWER ANCHOR INSTALLED 

SEP OCT NOV DEC JAN FEB MAR APR MAY JUN JUL AUG 

6.25 I I I I I I I I I I 

INSTRUMENT NO. SI - 66 - 3 

---m---y 

DEPTH = 25’ 

-0.50 - 

DEPTH = 0’ 

-“.75 - UPPER ANCHOR 

INSTALLED 

1982 1903 

LOWER ANCHOR INSTALLED 

-1 .oo I I I I I I I I I 

SEP OCT NOV DEC JAN FEB MAR APR MAY JUN JUL AUG 

NOTE: (+)MOVEMENT AWAY FROM EXCAVATION 

(-)MOVEMENT TOWARD EXCAVATION 

FIGURE 40 - INCLINOMETER DEFLECTION ‘JERSUS TIME 

l-A-45 
+1.0- 

+o.so- 

-0.5- SI-51-1 

+1.0 1 

-0.5 -I 

-1.0 1 

SI-5 l-2 

-0.5 

i 

51-51-3 

-1 .o 

UPPER ANCHOR 

INSTALLED 

LOWER ANCHOR INSTALLED 

I I I I 

SEP ’ OCT ’ NOV ’ DEC 

I I 

JAN ’ FEB ’ MAR ’ APR ’ MAY ’ JUN 

1982 1983 

LEGEND: 

0 - INCLINOMETER DATA 

A- SURVEY DATA 

NOTE: (+)MOVEMENT AWAY FROM EXCAVATION 

(-)MOVEMENT TOWARD EXCAVATION 

FIGURE 41 - COMPARISON OF INCLINOMETER TOP MOVEMENT SURVEY 

DATA VERSUS INCLINOMETER DATA 

l-A-46 
; 

0 

ii 

P 

-0.5- 51-66-2 

-0.5 Sl-66-3 

-1 .o 

UPPER ANCHOR LOWER ANCHOR INSTALLED 

INSTALLED 

I I I 

I SEP ’ OCT ’ NOV ’ DEC JAN ’ 

I I I I 

FEB ’ MAR ’ APR ’ MAY ’ JUN ’ 

1982 1983 

LEGEND: 

O-INCLINOMETER DATA 

A-SURVEY DATA 

NOTE: (+)MOVEMENT AWAY FROM EXCAVATION 

(-)MOVEMENT TOWARD EXCAVATION 

FIGURE 42 . COMPARISON OF INCLINOMETER TOP MOVEMENT 

S!JRVEY DATA ‘JERSUS INSLINOMETER DATA 

l-A-47 
,077 __- 

* .OI m.06 -.“” WILLIAMS R-&AD& 

4TLANTA 

-- --. _ t 

MOTEL 

HORIZONTAL DEFLECTION, FEET 

(ALL DEFLECTIONS ARE OUTWARD) 

NO SCALE 

IL .“l 

_1 

’ ATLANTA COCA COLA 

a: BOTTLING COMPANY 

up: / 

1 

- 

INTHOUSE I 

MOTEL 

VERTICAL DEFLECTION, FEET 

!ALL DEFLECTIONS ARE DOWN EXCEPT AS NOTEE WITH -) 

NO SCALE 

LEGEND 

b PILE MONITOR (6 FEET BELOW TOP OF WALL) 

A GROUND MONITOR OR BUILDING MONITOR 

0 SLOPE INDICATOR MONITOR 

FIGURE 43 . SUMMARY OF MAXIMUM HORIZONTAL AND VERTICAL 

SURi/EY MEAFUREMENTS 

1-A-46 
HORIZONTAL DEFLECTION 

- I I - I 

0 10 20 30 40 

WALL HEIGHT, FT. 

LEGEND 

o- FACE OF SOLDIER PILE 6 FEET BELOW 

TOP OF WALL 

n - - - GROUND MONITOR 10 FEET BEHIND 

FACE OF WALL 

A - .- GROUND MONITOR 30 FEET BEHIND 

FACE OF WALL. 

.lOVERTICAL DEFLECTION 

0 

D 

0 / 

.05 - 0 A 

0 Y 

/ 

0 I n - I 

0 10 20 30 40 

WALL HEIGHT, FT. 

FIGURE 44-HORIZONTAL AND VERTICAL SURVEY DEFLECTIONS 

VERSUS WALL HEIGHT 

l-A-49 
+.ia 

+.10 1 PILE 51 , LOWER ANCHOR 

LOCK-OFF DATE l/25/63 

PILE 51 . UPPER ANCHOR 

LOCK-OFF DATE : 1216162 

10 IO0 

TIME , DAYS 

ROD-SHORT ROD 

1 IllI 

1000 

10 lb0 300 1000 

TIME , DAYS 

FIGURE 45 

LONG TERM RELATIVE DEFLECTION IN BONDED ZONE 

l-A-50 
+.10 1 

PfLE 66 , UPPER ANCHOR 

LOCK-OFF DATE : 12/15/82 ROD-SHORT ROD 

5 

I 

“‘\ 

I I l!llll I I I r,,,, 

IO 

F J 

100 

300 

IO00 

iii 

i 

“‘\ 

‘k-.-’ 

A.-’ 

/*-* TIME , DAYS 

LONG ROD-SHORT ROD 

+.IO - PILE 66 . LOWER ANCHOR 

I/! 

w 

LOCK-OFF DATE : z/8/83 

3 

Z 

w 

u +.05 - 

G 

E 

it 

0 

5 O I 

F 

I I),,,, I I , I,,,, I I I I,,,, 

w” 

IO JO0 300 

too0 

L’ TIME , DAYS 

W 

n 

-.05 

FIGURE 46 

LONG TERM RELATI’dE DEFLECTION IN BONDED ZONE 

1-A-51 

APPENDIX B 

File #80-25 

November 5, 1981 

Mr. Kenneth P. Akin, Jr. 

Law Engineering Testing Company 

401 West Peachtree Street, N.E. 

Suite 1840 

Atlanta, GA 30308 

Re: Fulton County Permanent Tieback Wall 

Dear Ken: 

This letter will summarize my recommendations for instrumenting 

the above tieback wall. 

1. SUMMARY OF MBASUREiMENT METHODS 

The primary parameters of interest in evaluating performance are 

given in Table 1, together with desirable accuracies, recommended 

instruments, and probable accuracies. Details are discussed in 

turn below. The two "primary stations" referred to in the table 

will, as agreed during our October 21 meeting, be selected by you 

and Georgia DOT. 

2. HORIZONTAL DEFORMATION OF WALL FACE 

Inclinometers 

At the two primary stations, pipe will be tack welded to the full 

length of the soldier pile (see Figure l), and plugged at its 

lower end with a wooden plug. The top of the pipe should be at 

final grade, i.e., above the soldier pile top. The pipe will be 

set in place with the soldier pile and, after the porous concrete 

has set, Georgia DOT will drill a minimum 3-inch diameter hole, 

through the pipe, to about lo-ft below the bottom of the soldier 

pile. The actual depth will be selected when geologic cross 

sections at the primary stations have been drawn. The 

inclinometer casing will then be installed by grouting through a 

pipe inside the inclinometer casing, mating with a check valve in 

the bottom cap. This method is necessitated by space 

considerations, and the need to minimize the size of the steel 

pipe from the structural reinforcement standpoint. 

The minimum i.d. for the steel pipe is 3-l/2", and Georgia DOT 

should confirm that they have drill tools that will drill a hole 

no less than 3" diameter through a 3-l/2" pipe. The new "CPI" 

inclinometer couplings, discussed during our October 21st meeting, 

l-B-1 
should not be used as they protrude outside the casing o.d. and 

aggravate the space problem. Presumably Nicholson Construction 

Co. (NCC) will supply and install the pipe. 

Table 2 gives a listing of required materials, on the assumption 

they will be ordered from Slope Indicator Company. 

EDM Survevs on Wall Tarqets 

Presumably all will be handled by Georgia DOT. At each 

measurement station targets should be set at each anchor head 

level and at the wall top. Stations should include the two 

primary stations, the four adjacent soldier piles (locations of 

telltale load cells), and approximately 8 others, 

along the wall. 

equally spaced 

Arrangements will have to be made for transposing 

targets from soldier piles to wall face, 

targets before the wall is poured. 

presumably by setting out 

Conventional Survey on TOP of Wall, Usina Offsets from a Line of 

Transit 

This is a duplication of EDM measurements on to the wall top, and 

provides redundancy of a significant parameter. Reference points 

will need to be set well behind any zone of possible movement, and 

a line of transit established from these points. The ten 

measurement'stations should be the same as those used for EDM 

measurements, plus the location of the four soldier piles adjacent 

to the primary stations. Data will also provide a check on 

inclinometer data in the two casings at soldier piles. 

3. ABSOLUTE DEFORMATION OF ANCHORS 

Absolute deformation of anchors at each of the two primary 

stations will be measured by a combination of inclinometer and 

telltale readings. Deformation of the soldier pile at each anchor 

head will be determined using the inclinometer, to which will be 

added 

each 

the measured relative deformations between soldier pile and 

end of the bond zone. Relative deformations will be 

determined from two rod telltales installed with each anchor, one 

attached to a strand at the top of the bond zone, the other to the 

same strand at the bottom of the bond zone. 

each anchor will be the 

The upper telltale on 

same as used for the "telltale load 

cells," described later and shown on Fig. 5. Hence these serve 

both as deformation and load indicators. 

4. SETTLEMENT OF WALL 

Presumably all will be 

conventional 

handled by Georgia DOT, 

optical leveling procedures. 

using EDM or 

Targets should be at 

the same ten stations used for survey measurements on the wall 

face pulse the location of the four soldier piles adjacent to the 

primary stations. 

l-B-2 
5. HORIZONTAL DEFORMATION AND SETTLEMENT OF SURFACE 

OF SOIL BEHIND WALL 

Presumably all will be handled by Georgia DOT, 

conventional optical survey procedures. 

using EDM or 

Targets should be at the 

same ten stations used for survey measurements on the wall face 

(but no need for measurements at the locations of the four soldier 

piles adjacent to the primary stations). 

ft, 30 ft, 

Targets should be set 10 

and 60 ft behind the wall, and should consist of a 

sleeved rod or pipe anchored below the zone of seasonal vertical 

deformation. Figure 2 shows 

which would be suitable. 

the standard New York DOT detail, 

At four of 

inclinometer 

the locations the target will be the top of 

casings, described below. At these locations, the 

horizontal deformation measurements will provide a check on 

inclinometer data. 

6. HORIZONTAL DEFORMATION OF SOIL 

BEHIND WALL, BELOW SURFACE 

Inclinometer casings 

stations, 

will be installed at the two primary 

10 ft and 50 ft behind the wall. 

will be the 

Installation procedure 

injected down a 

standard Georgia DOT method, using a grout backfill 

tremie pipe outside the inclinometer casing. 

Table 2 indicates required materials, assuming the use of "CPI" 

casing. These should be ordered as soon as 

installations 

possible so that 

can be made in 

additional excavation at the site. 

the near future, before any 

In planning the installation 

method, note the coupling o.d. is 3.07". Top arrangements should 

be designed so that access will eventually be through a removable 

cover flush with the ground surface. 

should be. made to 

Comprehensive arrangements 

protect the installations 

construction equipment. 

from damage by 

7. LOAD IN ANCHOR STRESSING LENGTH 

The primary method of load measurement will be use of load cells. 

Telltale load cells, 

be used. 

as shown schematically in Figure 3, will also 

Load Cells 

A load cell will be installed on each of the two anchors, at the 

two primary stations. As discussed under "Load in Anchor Zone" 

below, a larger anchor 

anchors, 

head will be required for these four 

with an o.d. on the strand pattern of 4.65". Hence the 

load cell i.d. needs to be about 5". 

criteria are: 

Other load cell design 

0 225 kip capacity (to be confirmed when primary station are 

selected). 

l-B-3 
0 Adequate wall thickness for structural stability. 

o length: o.d. ratio not less than 1.4, to minimize errors due 

to end effects, i.e., a length of about 9". 

0 Rounded ends, to minimize errors due to end effects. 

0 Strain concentration at the gage points to maximize 

sensitivity. This can be accomplished by installing strain 

gages in side holes as on Figure 4. 

0 Hermetically sealing strain gages, and verifying seal 

integrity using a helium leak test. 

0 Use of a spherical bearing between bearing plate and load 

cell, to accommodate inevitable imperfect alignment. 

o Use of bearing washers above and below the load cell, with 

Rockwell hardness less than C32. 

o A waterproof connector, with dust cap, facing sideways. 

I know of only one company capable of supplying cells conforming 

to the above criteria: Brewer Engineering Laboratories, Inc. 

(BEL), Marion, Ma. They have many years of experience in field 

strain gage work, and have helium leak test equipment. This test 

provides a check on gage hermetic sealing, and essential feature 

when using resistance strain gages for long term applications. I 

met with BEL on October 28 to discuss our requirements, and have 

received from them the budget in Table 3. 

While at BEL, I discussed the need for providing a method of 

removing cells for check calibration at any time. In their 

experience,, 

there is 

given high quality hermetically sealed strain gages, 

Fry minor likelihood of a significant change in gage 

calibration over a 15 year period. However, the gage zero can 

change and, as our required accuracy is very high, they 

recommended a procedure for checking at any time. 

believe it is 

They do not 

necessary to remove the cells, and are concerned 

that the split shim arrangement used previously by NCC may create 

non-uniform loading on the upper end of the cell. They recommend 

a lift-off method, using a pull-bar, allowing the cell zero 

reading to be reestablished at any time. By strain gaging the 

pull-bar or by using a supplementary load cell, the calibration 

can also be verified during the lift-off tests. I support this 

recommendation. Details of this arrangement are best worked out 

by NCC with input from BEL as necessary. 

Use of load cells will extend the anchor head length by about 15 

inches (cell 9", 

2-l/21, 

two bearing washers l/4" each, spherical bearing 

additional bearing plate 2-3/4"). As discussed during our 

October 21 meeting, the soldier beams at the two primary stations 

l-B-4 
should be installed out-ot-line, further from the wall face, to 

accommodate the cells. A detail drawing will be required from 

NCC . 

Telltale Load Cells 

The principle of a telltale load cell is shown in Fig. 3. 

Assuming no creep in the strands over the stressing length, 

telltale load cells will provide backup to the load cell 

measurements. However, if any creep does occur in the stands over 

the stressing length, comparison between load cell and telltale 

load cell data will provide data on creep magnitude. 

As described earlier under "Absolute Deformation of Anchors" these 

telltales will also be used as deformation indicators. A 

schematic of rod telltale arrangements, both for local and 

deformation measurements, is shown on Figure 5. Note the location 

of the telltales at the anchor heads, the position with respect to 

tendon spacers, and the need for tendon spacers in the stressing 

length. 

Each telltale will consist of a flush-coupled l/4 inch stainless 

steel rod within an oil-filled flush-coupled l/4" Schedule BO PVC 

pipe (0.540" o.d., 0.302" i.d.). A disconnect will be provided 

between telltale anchor and telltale rod, so that telltale rods 

can be installed after anchor grouting and so that free-sliding 

can be checked at any time. Telltale anchors will be attached to 

strands using clamps and epoxy. Stainless steel reference plates 

will be required on the anchor heads. 

Installation will generally be as follows: 

o Attach telltale anchor and PVC pipe to strand, feeding 

pipe through appropriate tendon spacers. 

o Cement cap on upper end of PVC pipe. 

o Insert strands, pipe, etc., into corrugated plastic pipe. 

o Install and grout anchor. 

0 Cut off pipe cap. 

o Fill pipe with oil. 

o Insert telltale rod in pipe and lock in place. 

o Conduct anchor proof test and calibrate telltale 

"change in stickout" versus load, using the BEL 

load cell. 

Note that the final step will require that the last of the 4 

primarv station anchors must be proof tested after all 8 adjacent 

anchors, so that a BEL load cell can be used for all in-situ 

calibrations'. 

l-B-5 
Although the telltale principle is very straightforward, many 

details need to be finalized before components are machined. In 

my view, the expedient and efficient way if for me to coordinate 

directly with Geokon Inc., Lebanon, NH ( Dr. Barrie Sellers) to 

finalize details and prepare shop drawings. Components can then 

readily be machined by Geokon. Clearly there are alternative 

sources for these materials, but none allowing such close 

coordination. I have used Geokon for similar work previously, and 

have found them to be both efficient and competitive. Budget 

information is included in Table 4. Note the item for engineering 

time: This is for Barrie Sellers' time to coordinate with me and 

prepare shop drawings. 

8. LOAD IN ANCHOR BOND ZONE 

In my October 9, 1981, letter, I identified three possible methods 

of measuring load in the bond zone: resistance strain gages, 

vibrating wire strain gages, and telltales. Having studied these 

options 'further, I now consider that strain gages are not 

practicable, for the following reasons: 

0 At 225 kips the strands are subjected to about 7000 microstrain. This is nearly three times the range of the 

miniature vibrating wire strain gages and very large when 

considering long-term drift-free performance of resistance 

strain 'gages. 

0 The gage carrier plate must not reinforce the strand 

significantly. Thus the carrier plate cannot be robust, and 

would be very subject to damage while installing the anchor. 

o A proven method of clamping the carrier plate to the strand 

is not available. I've evaluated various options, but all 

would entail a significant test effort before they could be 

relied upon. 

0 There is a likelihood of torsional strains in the strands as 

they are tensioned, hence an uncertainty in converting 

measured strain to axial load. 

The alternative of multiple telltales is both practicable and 

economical. The principle is essentially the same as the telltale 

load cell, using any pair of telltales to create a gaged length of 

the bond zone. The system would be installed on the four anchors 

at the two primary stations. Sketches of the arrangement are 

shown on Figs. 6 and 7. 

Proposed details are: 

o A telltale anchor attached to one strand at 5 points in the 

bond zone, i.e., at 7' -6" spacing. The anchor surface will 

be knurled to match the strand outside irregularities, for 

good bond. 

l-B-6 
0 A wire attached to each anchor. 

0 A l/4" oil-filled nylon tube around each wire, connected to a 

tube fitting threaded into the anchor. 

o A 3/4" PVC pipe on the axis of the entire 45 ft anchor, with 

slots for grout entry and nylon tube entry, through which all 

nylon tubes and wires will pass to a central hole in the 

anchor head. 

o A small collar attached to the upper end of each wire, above 

the anchor head. 

o A chair for the indicator, attached to the anchor head. 

o A mechanical indicator, as shown in Fig. 8. 

The system is available from University of Newcastle Tyne, 

England. I've used the Mark 2 version on two projects requiring 

long wires, but for this project the Mark 1A (not the reversed 

version) would be used. A major attraction of this device is the 

ability to make a deformation reading on an individual wire at 

more than brie standard wire tension, because the difference 

between the' two readings should always be equal to the elastic 

elongation of the wire, i.e., always a known magnitude and always 

the same. Hence accuracy can be increased by reading at various 

tensions and drawing a straight line through the plotted points, 

and any lack of free wire movement will immediately be discerned. 

Installation will generally be as follows: 

0 

0 

0 

0 

0 

Build the anchor around the 3/4" PVC pipe, with appropriate 

bands and spacers as on Fig. 6. Note that this creates two 

non-standard features: spacers in the stressing length and a 

separation between strands at band points in the bond zone. 

I've checked with NCC whether these features are allowable, 

and they see no problem. Cut additional slots in the pipe as 

necessary. 

Pre-assemble anchor/wire/tube/oil assemblies with upper end 

of tube plugged internally. This will entail use of a wire 

straightener and oil pump, and an excess length of tube to 

house the full wire length. 

Feed nylon tubes into the 3/4" pipe, from the anchor 

locations. 

Attach telltale anchors to a strand. Attach a small block of 

Styrofoam above each anchor so that, if the strand pulls 

outward with respect to the grout during stressing, the 

anchor clamps will not slip. 

If necessary (to be determined during detail design), 

pressurize oil in the tubes by connecting all 5 tubes to a 

l-B-7 
common hydraulic line and locking off, to counterbalance 

subsequent grout pressure. 

0 Insert strands, etc., into corrugated plastic pipe. 

0 Install and grout anchor. 

o Assemble indicator chair, wire collars, attach indicator, and 

conduct proof test. 

Wire telltale readings during proof testing, will give data for 

calculation of stresses in the bond zone, using either the 

theoretical strand modulus or the modulus determined for the 

telltale load cell. A significant effort needs to be made to 

ensure compatibility between proof testing and telltale indicator 

arrangements at the head. Note that during stressing, it may be 

necessary to use more than one collar on each wire, as the 

indicator range is 2". Theoretically, 30 ft of 5 strand anchor 

loaded to 225 kips elongates about 2-l/2". 

As for the rod telltales, many details need to be finalized, and I 

recommend working with Geokon, Inc. to do this. The indicator and 

associated items can be procured directly by Georgia DOT from 

England, but I believe it would be more convenient to procure 

through Geokon, hence ensuring compatibility of all components. 

Budget information is included in Table 5. 

9. TELLTALE DATA REDUNDANCY 

The shortest and longest wire telltale should provide the same 

data as the two rod telltales (see Fig. 6). 

The load at the top of the bond zone will be known from load cell 

and telltale load cell data. The load at the bottom of the bond 

zone is, of course, zero. Hence the load transfer curve, plotted 

from wire telltale data, should be consistent with these two known 

load values. 

10. BUDGET FOR MATERIALS 

The materials budget is summarized in Table 6. The $33,075 amount 

includes a 10% contingency and also $1750 of engineering time by 

Geokon, Inc. 

11. MISCELLANEOUS 

(a) A stable benchmark will be required for survey 

measurements. 

(b) A suitable blockout needs to be made through the wall at 

the location of each instrumented anchor. Figs. 9 and 

l-B-8 
12. 

10 show approximate dimensions. Presumably each 

blockout will be formed with a steel box and, as 

discussed during our October 21 meeting, special care 

needs to be taken to ensure that the covers do not allow 

water to stain the face of the wall. We need to 

determine who will design, provide and install the steel 

boxes and covers. NCC should check that I've left 

enough room on Figure 9 for the lift-off arrangements. 

(c) We need to determine the access needs and arrangements 

for reading at the 12 instrumented anchor heads, both 

during and after construction. 

ACTION ITEMS BY GEORGIA DOT 

(a) 

w 

(c) 

w 

(e) 

(f) 

Review the recommendations in this letter. 

Confirm suitability of 3-l/2" min, i.d. for steel pipe 

on soldier piles. 

Order casing, etc., after confirmation of hole depths. 

Confirm availability of nearby stable benchmark. 

Authorize me to submit a load cell specification to BEL 

and to request that they submit a quotation to Georgia 

DOT. If acceptable, an order should be place quickly. 

Authorize Geokon, Inc. to proceed with preparing detail 

designs and shop drawings for rod and wire telltales, 

and to order appropriate items from England as on Table 

5 

13. ACTION ITEMS BY LETCO 

(a) Review the recommendations in this letter. 

(b) Determine required lengths of inclinometer casings. 

(c) Confirm load cell capacity is 225 kips (depends on 

selected primary stations). 

14. ACTION ITEMS BY NICHOLSON CONSTRUCTION CO. 

These assume acceptance of Nicholson's drawings, and on resolution 

of any contractual matters resulting from these instrumentation 

plans. 

(a) Make a thorough check on my proposed rod a;za;;;; 

telltale arrangements, installation and . 

procedures, to be certain that all is practicable. 

l-B-9 
w 

P-3 

w 

(e) 

w 

(9) 

w 

(i) 

Select and supply suitable pipe with wooden plug for 

inclinometers on two soldier beams. Minimum i.d., 3-l/ 

2" . If possible, thinner wall than standard pipe 

(standard 3-l/2" is 0.226" wall). Include on 

construction drawings. 

Prepare detail drawing showing primary station soldier 

piles installed out-of-line, with load cell, bearing 

washers, spherical bearing and additional bearing plate. 

Coordinate with BEL as necessary. Supply 4 additional 

bearing plates. 

Prepare detail drawing showing load cell lift-off 

arrangements. We need to discuss: 

o How will lift-off tests be made? Who will make them? 

Who provides what materials? 

o Whether to use a supplementary load cell or a strain 

gaged pull bar. I favor the latter, and recommend 

that NCC provide a pull bar (and end attachments for 

calibration) to BEL for strain gaging. 

Provide 48 additional tendon spacers for anchor 

stressing lengths to accommodate rod and wire telltales: 

o 4 in each of 4 anchors at primary stations. 

o 4 in each of 8 anchors adjacent to primary stations. 

Provide 4 anchor heads at primary stations with 9 holes. 

I understand these are standard, with 2-l/8" diameter 

threaded central hole and 8 holes 0.7" diameter, with 

head o.d. 5.95" and o.d. on strand pattern 4.65". Drill 

and tap heads to receive rod telltale reference surfaces 

and wire telltale indicator chair. Confirm you will 

have a jack with adequate center hole diameter for this 

pattern. 

Provide 5" pipe sleeve (instead of 4) at 4 anchors at 

primary stations. 

Provide 8 anchor heads at stations adjacent to primary 

stations, with approx. l-1/8" diameter central hole and 

6 outer 0.7" diameter holes (5 for strands, one for 

secondary grouting). o.d. on strand pattern approx. 

3.7". Drill and tap heads to receive telltale load cell 

reference surface. 

Assist with ensuring compatibility between wire telltale 

indicator and proof testing arrangements at the head so 

that deformation readings can be taken proof testing. 

1-B-10 
(j) Check adequacy on Figure 9 of room for lift-off 

arrangements. 

As discussed with you, I believe that all recipients of this 

letter should meet in the near future to resolve outstanding 

points so that we can proceed with critical items. NCC believes 

that the first anchor could be installed as early as mid January, 

and load cell delivery may be 8 weeks. A good date for that 

meeting would be Thursday November 12. According to my present 

schedule, I will be out of the country from November 20 through 

December 15, hence my availability is very limited. 

Sincerely, 

John Dunnicliff 

l-B-11 
Par-W 

kSlrrbl8 

kcurrcy 

Probrbl. 

kcurrcy 

norirontrl detorrtlon pt nil 

tree. 

? 0.1. fnclinaeter rtteched 

ta soldier pile at 

? 0.1' 

tw primry stations. 

Em'sumy on to 41 to be .usessed 

f9@ti@i&WP~. by 6eorg18 DOT. 

conmtiona; sumy 

wuuring ottseU t&m 

2 0.1‘ 

1 line of trsnsit, on 

tokrgettrttopet 

wll. 

Absolute detomatlon of 8nchor. ? 0.1. fellt8les on rrchon 

att*o prlmry statIons 

? 0.1. 

(8180 wing d8t8 tras 

inclinaeter rtt8ch8d 

to Soldier beu). 

frttlamnt of wll. 2 0.1' Conventionalsurveyor 

z z g1y9- at 

2 0.1. 

. 

Horltontrl drtonr- + 0.1' 

tfon and settlrmt - 

Canvention81 survey or 

Em! on to ShllGl 

? 0.1' 

of sulv8a of soil 

behind wall. 

reterwKe pollItS in 

mlol,~8ppmx. 10 

. 

HorIzontat detorw- ? 0.1' Imlinart8n 8ttro 

t1on of roll behind prirry statlaw 

? 0.1. 

wall* below Surf8ce. 

Lord in rnchor 

stressing length. 

Lo8d cells on abors 

rt t# priury stst1on8, with rrrrngemts for in-situ 

mrlibnt1m (4 

cells). 

lellt8le load cells 

onrnchonrt&o 

primry stationS 8ad 

at tour sdjscent 

secondely ststlons 

(12 tellt8leS). 

~O.WOf ra 

T 

* 

ipz!J hJ i&i. 

B 

alibrrtion 

rllous check on 

drift and ulibrrtion any tim. 

t2toSkipr 

it no creep In 

strmdr. ti 

pertson with 

lord cell gives 

drta on 8ny 

creep in strrnds 

over stressing 

length. 

Lord In anchor 

bond mm. 

As for Ilu1ti-uire tellt8les t2 to 5 tips stressing ori8nchors8t&o length. primry ststlons, 

~~y~reep In 

. 

c!&iiiul . 

l-B-12 
TWE 2 

1NCl1MKlTREQUIPMENTTDBE@DERED~ 

SLOPE 1RDIcAwR CowAm 

Part I 

wt 

Oescrlptlon Ourntj tu price kokmt 

For Caslnm on Soldhr Piles 

51111 2.75' o.d. plrrtlc using 

In 10 ft lengths 

51112 2.75' o.d. plastic 

couplings 

51115 Protectlve top up 

51133-2 Pop rivets 

51134-l klvent.cnt, l/2 pt. 

51136 Cmut plug, gasket 

seal type, for 2.75' 

plrttlc using 

For Casinos In Soil 

57511 2.75' o.d. CPI aslng 

In 10 ft lengths 

57512 3.07 o.d. CPI couplings 

57515 CPI bottm up 

51115 ProtectlVr top up 

57517 sherr win 

57540 '0' ring lubricant 

TOTAL mwr 

"bl' 

10 

2 

100 

1 

2 

911' 

20 

4 

4 

so 

1 

$4.15 

1.50 

2.5s 

.12 

2.50 

62.00 

4.50 

7.50 

9.00 

2.w 

2s 

2.2s 

S 456.SD 

15.00 

5.10 

12.00 

2.50 

124.00 

990.00 

lSD.00 

%.DD 

10.20 

12.so 

2.25 

tl,816.D6 

1. AstImes use of LETCO owed Dlgltllt, able, lndlcrtor, pulley usealy 

and cable hold. 

2. ksus 6eorglr DDT rlreedy ems usdly tools for stmderd 2.75' 

o.d. plastic uslnq. 

3. ksus bottm of using Is 10 ft below botta of soldier pile end 

tImtone 10 ftlength 1s cut end used for the top length of eech 

lnstallrtlon. 

4. ksus 6eorglr #nut11 plm end provide top protective ruengemtt. 

1-B-13 
TABLE 3 

LOAO CELL EQUIPUEMT 

Recomnended Supplier: Brewer Engineering Laboratories, Inc. 

P.O. Box 288 

Marion, HA 02738 

(617) 748-0103 

AllNi Mr. Leon Ueyntouth (or) 

Mr. Verne Yallace 

Description Quantity 

Unit 

Price* AmJnt* 

Load cell, including calibrations, 2 bearing washers and 

spherical bearing plate 4' $2,690 $10,760 

Vishay P-35OA strain 

indicator with connector 

and 10 ft jumper cable 1 1,195 1,195 

Vishay gage installation 

tester 1300 1 695 695 

Gage and calibrate pull-bar 

supplied by others. 

Includes cable and connector 

but excludes any necessary 

end fittings on pull-bar 

required during calibration 

(i.e., assumes they will be 

provided with pull-bar) 

Instruction manual and data 

sheets L.S. 

1,900 

so0 

1,900 

500 

TOTAL AMOUM 

(excluding freight) $15,050 

*Prices for budgeting purposes only. These are not a fonrml quotation. 

Prior to obtaining a formal quotation a tpecificatfon needs to be 

written, as per the reccmendations in this letter. Note that delivery 

is 6 to 8 weeks, hence arrangements need to be finalized as soon as 

possible. 

1-B-14 
WlE 4 

(Do TEllIKE EQJIPSWl 

lmt 

OulntW Mu(S) 77’ 

l/4' stalnlrss steel flush 

coupled rod (wlous 

lengths) 

l/4' Sch. 60 WC pip 

In 2Dttleqths 

Ditto, u/flush thre,r&d 

touollnoI 

l/4' PVC pipe ups 

PVC cant, brushes 

Telltale anchor ulth mlr 

mnd femle disconnect, 

mnd clrrps 

Anchor he8d refermc8 

sutfrce for cmw hole, 

w/telltale mr 

Dltto for outside hole 

S.S. chrlr for In-situ 

ullbration 

Ditto for rerdlng 

nicrorter for In-sltu 

c811br8t1on 

Ditto for rwedlng 

. 

Dild~illing 

Installrtlon tools 

Illscellrneous mterlrls 

Englneerlng iir 

lDlKMDWl,EXCLUD166FRE16W 

36D tt 2.00 

240 tt 30 

16Dft 2.00 

16 1.u) 

1 Pt lO.DD 

16 

6 

6 

2 

2 

1 

1 

1.3. 

L.S. 

L.S. 

4D.w 

16.00 

15.00 

10.00 

10.00 

166.00 

166.00 

Included onleble 5 

lDD.DD 

260.00 

2arr 3SD.W 

720.00 

12D.w 

#o.DD 

24.00 

10.00 

‘Prices for budgeting pwposes Only. 

1-B-15 













REPORT NO. 2 

PERFORMANCE OF TWO 

INSTRUMENTED, PERMANENT POST-GROUTED TIEBACKS 

THE MT. CARMEL CEMETERY RETAINING WALL 

I-95 - BALTIMORE, MARYLAND 

bY 

Harald P. Ludwig 

Report submitted by Schnabel Foundation Company for: 

Permanent Ground Anchor Demonstration Project 

Federal-Aid Project No. I-95-4(118)57 

State Highway Administration No. BC 246-123-815 

Baltimore City No. 2641 

April 1984 

TABLE OF CONTENTS 

I. Introduction . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2-l 

II. Site Conditions .......................................... 2-l 

III. Tieback Description and Installation Procedure .......... 2-2 

A. Tieback Description .................................. 2-2 

B. Installation ......................................... 2-3 

IV. Instrumentation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2-4 

V. Test Program . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2-6 

A. Initial Test Program ................................. 2-6 

B. Two-year Monitoring Program .......................... 2-7 

VI. Results and Analysis ..................................... 2-8 

A. Short-term Investigation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2-8 

1. Tieback 15-8 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2-10 

a. Load-displacement-time Performance ........ 2-10 

b. Strain Distribution ....................... 2-12 

(1) Tendon .............................. 2-12 

(2) Deformed Metal Tube ................. 2-13 

c. Computed and Measured Movements . . . . . . . . . . . 2-14 

2. Tieback 18-2 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2-15 

a. Load-displacement-time Performance ........ 2-15 

b. Strain Distribution ....................... 2-15 

(1) Tendon .............................. 2-15 

(2) Deformed Metal Tube ................. 2-17 

c. Computed and Measured Movements . . . . . . . . . . . 2-17 

3. Behavior Comparison . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2-18 

B. Long-term Investigation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2-19 

VII. Conclusions . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2-21 

VIII. Recommendations . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2-22 

2-i 

MT. CARMEL CEMETERY 

DEMONSTRATION TIEBACKS 

I. Introduction 

In January and February of 1982, Schnabel Foundation Company 

installed and tested two instrumented, permanent post-grouted 

tiebacks (designated as tieback 15-8 and tieback 18-2) in a stiffto-very-stiff, red-brown, silty clay at Mt. Carmel Cemetery 

retaining wall located alongside Interstate 95, north of the 

Baltimore Harbor Tunnel in Maryland. The two instrumented 

tiebacks were to be incorporated into the overall plan of 

stabilizing the failed retaining wall with 98 tiebacks and 4439 

linear feet of drains. The purpose of this demonstration project 

was to study both the short-term and long-term behavior of 

permanent tiebacks anchored in cohesive soils. 

The short-term investigation consisted of conducting one-week long 

creep tests on each tieback in order to study their loaddisplacement-time behavior. The long-term investigation consisted 

of monitoring changes in tieback load, wall deflection, movement 

of the soil mass, and changes in water table level over a two-year 

period. The Maryland State Highway Administration Bureau of Soils 

and Foundations instrumented and monitored the retaining wall and 

soil mass, while Schnabel Foundation Company was responsible for 

the instrumentation and monitoring of the tiebacks. 

For purposes or organization, the remainder of this report is 

divided into seven sections. Site conditions are described in 

Section II. The tieback description and installation procedure 

are given in Section III. Instrumentation of the tieback is 

discussed in Section IV, while the test program is described in 

Section V. Results of the test program and analysis of the 

performance of both tiebacks form the contents of Section VI. 

Finally, conclusions and recommendations for further study are 

given in Section VII and Section VIII, respectively. 

II. Site Conditions 

The two instrumented tiebacks are located approximately at Station 

304+23 and Station 305+19. Cross sectional views of the retaining 

wall and instrumentation as well as soil profiles at these 

stations are shown in Figures 1 and 2. The soil profiles indicate 

that both tiebacks were installed in a stiff-to-very-stiff, redbrown, silty clay. Rebar spacing at the front face of the wall is 

12 inches both horizontally and vertically. Rebar spacing at the 

back face of the wall varies from 6 to 12 inches horizontally and 

is 12 inches vertically. 

2-l 
III. Tieback Description and Installation Procedure 

A. Tieback Description L 

It is prudent to mention some of the terminology that will be used 

in the ensuing description of the tieback and in the remainder of 

this study. Every tieback has an anchor length and an unbonded 

length. The anchor length is the designed length of the tieback 

where the tieback force is transferred to the soil. This part of 

the tieback is commonly referred to as the anchor. The unbonded 

length of the tieback is the length which is free to elongate 

elastically. Both tiebacks had an anchor length, la, of 30 feet 

and a minimum unbonded length, lu, of 27 feet. 

Post-grouted tiebacks, known as TMD ("terrain meuble deferment") 

tiebacks developed by SIF Bachy of France, were used in this 

demonstration project because: 

0 they provide corrosion protection for the lifetime of the 

structure; 

0 they are capable of developing high capacities in cohesive 

soils; 

0 they are capable of being instrumented; 

0 the installation procedure causes minimal damage to the 

instrumentation; 

0 they were to be installed in small diameter holes in the wall; 

and 

0 installation equipment did not require a large construction 

easement along the wall. 

The components of the TMD instrumented tiebacks are shown in 

Figure 3. The TMD tieback used in this study consisted of a 3- 

inch diameter deformed metal tube, which is grouted to the soil, a 

3-l/2-inch diameter polyvinyl chloride (PVC) pipe which served as 

the bond breaker, and a 60-feet long, 1 3/8-inch diameter threaded 

Dywidag bar. The inflatable bag shown in Figure 3 is optional and 

was not used in this study. Its primary purpose is to allow the 

post-grouting operation to commence immediately after the TMD 

assembly is inserted in the hole. The threaded Dywidag bar, which 

is inserted in the deformed metal tube, had heat-shrink tubing and 

polyethylene bond breaker over its unbonded length. 

2-2 
B. Installation 

The installation procedure consisted of four major operations (see 

Figure 4): 

o punch an 8-l/2-inch diameter hole through the wall; 

o drill hole through soil to desired depth and insert TMD 

assembly; 

o grout the deformed metal tube to the soil; and 

o grout tendon to the deformed metal tube. 

An 8-l/2 inch diameter hole was punched through the wall by using 

an 8-inch diameter down-the-hole hammer with a button bit attached 

to its end. Upon encountering the first section of rebar 2-l/2 

inches from the front face of the wall, a laborer used a torch to 

burn off the rebar. The second section of rebar, located near the 

back face of the wall, was burned out with a lance. 

Once the hole had been punched through the wall, the hole was 

drilled to a final depth of 57 feet using a 6-inch diameter auger 

with a clay bit attached to its end. From the experience gained 

during the installation of the 96 non-instrumented tiebacks, it 

had been decided the hole could be drilled uncased. This greatly 

simplified the drilling operation. Very soft, soupy soil with 

running sand was encountered while drilling the hole for tieback 

18-2. Fortunately, this presented no problems during the 

installation of the tieback. 

Once the auger had been extracted from the drill hole, the 

deformed metal tube, with grout valves located every 3.3 feet 

along its length, and the PVC pipe were placed in the hole. The 

assembly was centered in the hole by means of plastic centralizers 

attached to the deformed metal tube. After the TMD assembly had 

been inserted in the drill hole, a 3-inch diameter double packer 

was positioned inside the deformed metal tube opposite the bottom 

grout valve. The packer was inflated and grout subsequently was 

pumped through the last valve until the grout could be seen 

exiting from the drill hole. The packer was removed and the 

inside of the TMD assembly was flushed with water and blown out 

with compressed air until the effluent was fairly clear. The 

grout was then allowed to set up overnight. 

The post-grouting operation began on the next day. 

positioned inside 

The packer was 

the deformed metal tube opposite the bottom 

grout valve and inflated. Grout was pumped slowly through the 

valve until either a maximum grout pressure of 40 bars was reached 

or one bag of cement had been pumped through the valve. The 

quantity of grout pumped through each valve was limited, because 

2-3 
REPORT NO. 3 

PERMANENT TIEBACK ANCHORS IN COHESIVE SOIL 

DEMONSTRATION SITE 

SR-90 (SR-5 to CORWIN PLACE) 

SEATTLE, WASHINGTON 

bY 

Hart Crowser and Associates, Inc. 

WASHINGTON STATE 

DEPARTMENT OF TRANSPORTATION 

MARCH 1984 
0 a calibrated 150-ton capacity load cell to monitor load at the 

anchorhead during the short-term investigation; 

0 a calibrated test pump and jack to load the tieback; 

0 two dial gauges to monitor movement of the anchorhead relative 

to a fixed point to the nearest .OOl inch; 

0 a Vishay 220 digital strain readout unit capable of reading and 

recording the response of 30 strain gauges continuously at a 

rate of one strain gauge per second; 

0 a Vishay P-350A strain indicator to monitor load in the load 

cell; and 

0 a dummy gauge used to zero the Vishay 220 digital strain 

readout unit prior to each set of readings over the two-year 

monitoring period. 

In general, the strain gauges performed quite well during the 

initial two-week test program. Strain gauge failure occurred in 

less than 17 percent of the strain gauges. However, the 

performance of the same strain gauges over the two-year monitoring 

period was less satisfactory. This type of performance is to be 

expected from electrical resistance strain gauges which are very 

accurate but, due to electrical aging of components and creep in 

bonding agents, tend to "drift" with time. Prolonged exposure to 

a hostile environment over a period of time was probably the 

reason for the majority of the strain gauge failures. 

Since the original intention of the two-year monitoring program 

was to monitor changes in tieback load, the strain gauges of 

primary interest were those located in the unbonded length of each 

tieback. Fortunately, 8 out of the 10 strain gauges located in 

the unbonded length functioned properly for the entire two-year 

monitoring period. Therefore, the two-year load change data 

presented in Section VI is considered reliable and accurate 

provided that certain limitations associated with the use of 

strain gauges are recognized. First, the tendency of electrical 

resistance strain gauges to ttdriftl# II with time influences the 

absolute magnitude of load measured more so than the trend in load 

change. Both the strain gauges installed on the two tie backs and 

the dummy gauge used to zero the Vishay 220 unit were affected by 

this "drift" phenomenon. Second, the variation in Young's modulus 

of elasticity of the bar required to convert strain to load 

introduces error to the absolute magnitude of load measured. The 

conversion of strain to load required the use of Equation (1): 

P= eEA 

where P = load 

. ..(l) 

e= strain (x 10 -6 ) 

E = Young's modulus of elasticity 

A = cross-sectional area of the tendon 

2-5 
Figure 31. Measured Verius Computed Creep Movement of Tieback 18-2 

- c.. 

, c 

60 

33 

20 

0 

-w- Computed 

Measured 

0 0. 02 0. 04 0.06 0. 08 0. 10 

CREEP MOVEMENT (IN. > 

2-A-31 
In setting up each tieback test, extreme care was exercised to 

ensure that the bearing plates, load cell, and jack were aligned 

properly. This was done in order to minimize the effects of 

friction and eccentric loading due to mis-alignment. Prior to the 

start of each test, each strain gauge was calibrated internally 

and initialized to give a "zero" reading to which all subsequent 

strain gauge readings were referenced.' A 5.5-ton alignment load 

was then placed on the tieback to seat and align the testing 

hardware. At this time, the dial gauges were set up to measure 

movement of the anchorhead. All subsequent dial gauge 

measurements were referenced to the initial "zero" reading at 5.5 

tons. The test proceeded by setting the desired load on the 

Vishay P-35OA strain indicator. The jack was then pumped up to 

the desired load. At this time, the dial gauge and all strain 

gauges were read. If the next load was to be held constant, the 

jack was pumped up to the desired load within one minute. An 

initial reading was taken when the desired load was first reached. 

The one-minute reading of the load-hold corresponded to the time 

that had elapsed since the pump was first activated. That is, the 

time lapse between the initial '1 zero '1 reading and one-minute 

reading of the load-hold was always less than 60 seconds. During 

the load-hold, it was necessary to adjust the hydraulic pressure 

in the test jack in order to maintain the load constant. Load was 

decreased to the alignment load after each load-hold by setting 

the Vishay P-350A strain indicator to the alignment load and 

opening up the release valve on the pump. At the end of each 

test, the load was reduced to zero and all strain gauges were read 

and checked for internal calibration. Subsequently, each tieback 

was locked-off at 40 tons by reloading the tieback to 40 tons, 

tightening a nut against the bearing plate, and releasing the load 

from the jack. 

B. Two-year Monitoring Program 

The two-year monitoring program consisted of monitoring the change 

in load in each of the tiebacks at least once every three months 

during this period of time. The strain gauge readings taken 

immediately after each of the tiebacks was locked-off at 40 tons 

served as the initial zero to which all subsequent readings were 

compared. 

The procedure used to take a set of readings is described below. 

The dummy gauge was used to zero each of the 30 channels on the 

Vishay 220 readout unit prior to visiting the site. At the site, 

the dummy gauge was used again to re-zero the 30 channels. 

Subsequently, the strain gauges on both instrumented tiebacks were 

read within a period of 60 minutes. 

Note that the strain gauge readings taken in June of 1982 are 

omitted from this report because they were considered unreliable. 

The reason for their un-reliability was traced to the failure to 

pot the military connectors which served as the junction between 

the strain gauges and the Vishay 220 readout unit. Grouting of 

2-7 
Figure 29. Distribution of Residual Strain in the Deformed Metal Tube of 

Tieback 18-2. 

400 r * 13.9 Tons 

+ 25.4 

300 l 37.9 

v 

A 

50.8 

--. 

200 c >A 

4, 

0 63.3 

0 75.4 

I A’ \ / A 88.1 

-100 

-200 

-300 

-400 L ----l--- IL-~-----l------.-~L ~--L.-..---.- J I I 

0 5 10 15 20 25 30 35 40 

DISTANCE FROM BACK OF ANCHOR (FT. > 

-- ----. ------ .-.- - .-.- -- - -..- - --._--I-.--_ --._ - ____ _ -.. .._- .___ ____._ _ ._. - _.__..___ - _---_-- .-.. ---..-_-___ 

s*v-‘,s->s%s~~u~~I(“- ___l,l 
Residual anchor movement is a very useful measure of tieback 

performance, since it is a function of both the tieback 

construction technique and the soil in which the tieback is 

anchored. Reese (1976) highlighted the fact that the 

displacements required to mobilize ultimate capacity in bearing 

are much larger than the displacements required to mobilize 

ultimate capacity in shear in most soils. A large number of tests 

conducted on tiebacks using extensometers have verified that the 

anchor moves when loaded and partially rebounds when unloaded, and 

that the residual anchor movement can be measured reliably in a 

cyclic test such as the performance test (Schnabel, 1982). Since 

larger anchor displacements are required to mobilize greater 

tieback capacity (for loads less than ultimate), residual anchor 

movement tends to increase as the applied load increases. To 

date, no acceptance criteria exists to evaluate tieback 

performance based on residual anchor movement. 

Time-dependent tieback displacements at constant load are used to 

evaluate the long-term performance of a tieback. The timedependent or creep movements typically are plotted on a semilogarithmic scale as a function of time (see Figure 8). For each 

load, the creep movement per logarithmic cycle (or creep rate) is 

determined by computing the slope of a best fit straight line 

drawn through the data points. If a straight line cannot be 

drawn, the creep rate is determined by computing the maximum slope 

formed by the data points. Creep rate can then be plotted as a 

function of load. To date, tieback specialists cannot agree on 

the acceptance criteria to be used to evaluate tieback performance 

based on time-dependent movements. 

The preceding discussion has focused its attention on the 

measurements used by the engineering profession to evaluate a 

tieback's performance. Elastic movements, residual anchor 

movements, and time-dependent movements are used to assess both 

the short-term and long-term load carrying capability of a 

tieback. The ensuing discussion presents a step-by-step 

description of the analysis used herein to evaluate a tieback's 

performance based on computed performance which can provide 

insight into the measured performance. 

The most obvious contribution of the computed performance of each 

tieback is the determination of the distribution of strain, and 

therefore loadl, along the tieback for a given applied load and a 

given time (see Figure 9). The shaded area represents the total 

amount of strain that developed during a particular load-hold. 

The strain distribution in the unbonded length can verify easily 

the provision of a minimum desired unbonded length--a criteria 

used by the entire engineering profession to evaluate a tieback's 

performance. The installation of strain gauges at critical 

locations along the unbonded length, e.g., outside and inside the 

face of the tied back structure and at the transition from the 

unbonded length to the anchor length, can provide valuable 

information concerning the influence of friction and bending on 

the distribution of strain in the unbonded length. Typically the 

2-9 
l------I- l- .- 

0 

-- 

- --I__- 

--- 

, 

2-A-27 
Schnabel Foundation Company on instrumented straightshafted tiebacks anchored in clay have shown that 

tieback failure is usually associated with residual 

anchor movements greater than 0.5 inches. The loaddisplacement response of tieback 15-8 (see Figure 10) 

was much more elastic than that of the straight-shafted 

tieback (see Figure 7). Both the elastic loaddisplacement response and the relatively small residual 

anchor movements of tieback 15-8 indicate that 

additional capacity existed beyond 88.1 tons. 

The time-dependent response of tieback 15-8 ' 

presented in Figures 12 and 13. At the maximum tei: 

load of 88.1 tons, the creep rate was essentially 

constant at ,043 inches/logarithmic cycle in both the 

third logarithmic cycle (100 to 1,000 minutes) and the 

fourth logarithmic cycle (1,000 to 10,000 minutes). 

However, this creep rate was double the creep rate that 

occurred in the second logarithmic cycle (10 to 100 

minutes). The change in creep rate from the second to 

the third logarithmic cycle has a plausible 

explanation. In jacking the load up to the final test 

load of 88.1 tons, the load was inadvertently 

maintained constant for a period of four minutes at a 

load of 86.4 tons. Once the error was discovered, the 

load was immediately jacked up to 88.1 tons and 

thereafter maintained constant. It is apparent that 

the creep movement of the tieback at 86.4 tons, and the 

effect of the previous loads, influenced the creep 

performance of the tieback at a load of 88.1 tons by 

reducing the amount of creep that occurred in the first 

two logarithmic cycles. The influence of load history 

on the time-dependent performance of tiebacks anchored 

in cohesive soils is been noted in other tests on 

instrumented and non-instrumented tiebacks. 

The creep performance of tieback 15-8 showed no signs of 

imminent failure. The maximum creep rate of 0.043 inches/ 

logarithmic cycle is relatively small for a tieback 

anchored in cohesive soil. Even at the maximum test load 

of 88.1 tons, the creep behavior was essentially linear 

with respect to the logarithm of time after the second 

logarithmic cycle. Temperature changes can affect the dial 

gauge support, tendon, and hydraulic fluid in the test 

jack. Temperature-induced movements may become significant 

when anchor creep is small. Fluctuations in creep movement 

observed during the fourth logarithmic cycle at the 88.1 

ton load increment are attributable primarily to 

temperature changes between day and night. 

maintain the load constant at 88.1 tons, it 

While trying to 

was 

to adjust 

necessary 

the jack pressure in response to temperature 

changes. 

2-11 
.10 

. 08 

.06 

.04 

. 02 

0 

Figure 25. Creep Rate Versus Load for Tieback 18-2 

I 1 

0 10 20 30 40 50 60 70 80 90 100 

LOAil (TLINS> 
The most interesting aspect of the time-dependent 

distribution of strain in the anchor is the 

relatively uniform increase in strain at most 

strain gauge locations for one load. As a 

result of any this kind of behavior, the slope of the 

strain distribution curve between most strain gauge 

locations 

load-hold. 

remained essentially constant during the 

This implies that there was little 

time-dependent change in the rate of load transfer. 

Also note that the magnitude of 

strain increase 

time-dependent 

at the maximum test load of 88.1 

tons, which was held constant for 10,000 minutes, 

was not significantly greater than that of any 

other load-holds which were held constant for much 

shorter periods of time. 

Residual-strain distribution in the tendon is 

presented in Figure 15. 

residual 

In the unbonded length, 

strain decreased with increasing load. 

The reason for this behavior was 

hysteresis 

probably due to 

(see Section IV). The residual-strain 

distribution in the anchor conformed to expected 

behavior. As the load increased, the residual 

strain in the anchor increased. The residualstrain distribution indicates that the minimum 

unbonded length was provided. For all loads, the 

transition zone between the unbonded length and the 

anchor length was well-defined and occurred 27 to 

30 feet from the back of the anchor. 

Deformed Metal Tube 

The distribution of strain in the deformed metal 

tube is presented in Figure 16. The most 

interesting feature of the strain distribution in 

the deformed metal tube is the presence of 

compressive strains at the front of the anchor. 

This also indicates that the load applied at the 

anchorhead was transferred down the tendon to the 

front of the anchor and then transferred to the 

soil; otherwise, compressive strains could not 

exist at the front of the anchor. Also, note that 

the location of the neutral axis (transition 

from compression 

point 

to tension) gradually moved down 

the anchor as the load increased. 

The time-dependent distribution of strain in the deformed metal tube shown in Figure 16 differed 

significantly from that in the tendon 

14). During 

( see Figure 

the load-hold at 88.1 tons the deformed metal tube exhibited a significant 

increase in strain in the back of the anchor. This 

2-13 
90 

80 

60 

Figure 23. Residual Anchor Movement of Tieback 18-2 

I 

0 

2 40 

-I 

30 

20 

0 

0 0. OS 0. 10 0. 1s 0. 20 0. 2s 0.30 

RESIDUAL MOVEMENT <IN. ) 

2-A-23 
most loads. However, at the higher loads, the measured 

creep movements tended to deviate from the computed creep 

movements. It is highly unlikely that the back of the 

anchor had begun to move through the soil in light of the 

load-displacement-time performance and strain distribution 

characteristics of tieback 15-8. In addition, it should be 

noted that the maximum difference between the computed and 

measured creep movement is on the order of 0.06 inch. 

2. Tieback 18-2 

a. Load-displacement-time Performance 

Figure 22 presents a plot of load versus displacement 

measured at the anchorhead, while Figure 23 displays 

the residual anchor movement of tieback 18-2. At the 

maximum test load of 88.1 tons, the residual anchor 

movement was only 0.210 inches. Both the elastic loaddisplacement response and the relatively small residual 

anchor movements of tieback 18-2 indicate that it had 

additional capacity beyond 88.1 tons. 

The time-dependent response of tieback 18-2 presented 

in Figures 24 and 25 showed no signs of tieback 

failure. At the maximum test load of 88.1 tons, the 

creep rate was essentially constant at 0.020 inches/ 

logarithmic cycle for the first three logarithmic 

cycles (1 to 1,000 minutes). The creep rate could not 

be determined for the last logarithmic cycle, because 

of the significant decrease in movement throughout this 

cycle. This behavior is more likely attributable to a 

shift in the independent reference point from which 

anchorhead displacement was measured rather than an 

indication of actual physical performance (after the 

l,OOO-minute mark of the load-hold at 88.1 tons, the 

weather suddenly turned warm and caused the ice, which 

partially supported the independent reference point, to 

thaw). The apparent decrease in creep movement also 

masked the true residual anchor movement measured at 

88.1 tons. The true residual anchor movement was most 

likely somewhat greater than the measured 0.210 inches. 

b. Strain Distribution 

(1) Tendon 

The distribution of strain in the tendon of tieback 

18-2 is presented in Figure 26. There was very 

little decrease in strain along the unbonded length 

at most loads. Although bending strains were 

measured in the unbonded length, paired strain 

gauges were used to compensate for them. The 

magnitude of strain measured at the transition from 

the unbonded to the anchor length is interesting in 

2-15 
Figure 21. Measured Versus Computed Creep Movement of Tieback 15-8 

70 

60 

Gj 50 

El 

t 

30 

0 

-mm Computed 

Measured 

0 0.02 0.04 0.06 0. 08 0. 10 

CREEP MOVEMENT (IN. > 

2-A-21 
The time-dependent distribution of strain in the 

anchor shows a relatively uniform increase in 

strain at most gauge locations--behavior also 

exhibited in tieback 15-8. This implies that there 

was little time-dependent change in the rate of 

load transfer. The magnitude of strain increase at 

the maximum test load of 88.1 tons, which was held 

constant for 10,000 minutes, was not significantly 

greater than that of any other load-holds which 

were held constant for much shorter periods of 

time. 

Residual-strain distribution in the tendon is 

presented in Figure 27. In the unbonded length, 

residual strain decreased with increasing load, as 

was the case for tieback 15-8. The residual-strain 

distribution in the anchor length conformed to 

expected behavior. As the load increased, the 

residual strain in the tendon increased. The 

transition zone between the unbonded length and the 

anchor length was well-defined and occurred 28 to 

30 feet from the back of the anchor. This 

indicates that the minimum unbonded length was 

provided. 

(2) Deformed Metal Tube 

The distribution of strain in the deformed metal 

tube is presented in Figure 28. Strains tended to 

decrease to zero at the front of the anchor. The 

time-dependent distribution of strain in the 

deformed metal tube shown in Figure 28 exhibited a 

rather uniform increase in strain at each load at 

most gauge locations. In addition, very little 

load was transferred to the back of the anchor, 

even at 88.1 tons. Residual-strain distribution in 

the deformed metal tube is shown in Figure 29. 

Significant compressive residual strains were 

measured at the front of the anchor. The 

compressive residual strains increased as the load 

increased. Very little residual strain was 

measured in the back of the anchor for loads less 

than 88.1 tons. 

C. Computed and Measured Movements 

In the ensuing discussion of computed and measured 

movements, it should be noted that the computed 

movements of the anchorhead would tend to be less than 

the measured movements for reasons identical to those 

presented in the analysis of tieback 15-8 (see Section 

VI.A.1.c). At 88.1 tons, the measured total movement 

and creep movement at the l,OOO-minute mark were used 

because of the shift of the reference point. A 

2-17 
X 

2-A-19 
After the second logarithmic cycle of the load-hold at 88.1 

tons, the creep rate of tieback 15-8 was double that of tieback 

18-2. The greater creep rate exhibited by tieback 15-8 

appeared to be reflected by the time-dependent changes in 

strain along the deformed metal tube (see Figures 16-18). 

A comparison of the distribution of strain in the tendon of 

each tieback reveals that, for any given load, tieback 15-8 

exhibited slightly higher strains in the unbonded length. This 

behavior is most likely attributable to the presence of greater 

friction forces and bending strains in tieback 18-2. Both 

factors would tend to decrease the magnitude of tensile strain 

measured in the unbonded length. In the anchor length. The 

shape of the strain distribution curves of each tieback is 

different; however, relatively little strain was measured at 

the back of either tendon, even at the maximum test load of 

88.1 tons. The discussion on the mobilization of maximum load 

transfer indicated that a maximum value of load transfer had 

not been developed beyond the mid-point of the anchor length of 

either tendon. Based on the assumption that at failure a 

maximum value of load transfer is developed at the very back of 

the anchor, it appears that both tiebacks had additional 

capacity. A comparison of the distribution of residual strain 

in the deformed metal tube of each tieback reveals that much 

more residual strain was developed at the back of the anchor of 

tieback 15-8. 

B. Long-term Investigation 

The results of the long-term investigation of the two instrumented 

tiebacks are presented together with the long-term performance of 

the retaining 

the tiebacks, 

wall and soil mass to show that the performance of 

and the performance of the retaining wall and soil 

mass are interrelated. 

Figure 32 presents data related to the drop in water table level, 

wall movement, and tieback load loss at Station 304+23 and Station 

305+19. At Station 304+23, the data shows that the major drop in 

water table level was accompanied by a significant portion of the 

total wall movement back into the hill in the direction of the 

cemetery. In turn, the largest loss in tieback load occurred 

during the same time period. By November of 1982, the water table 

level has stabilized at an elevation 14 feet lower than its 

original position. After November of 1982, the wall movement 

appeared to be at least partially cyclical in while the 

tieback 

nature, 

load loss appeared to have stabilized. A maximum loss in 

tieback load of 9.2 tons was recorded in December of 1983. At the 

elevation of the tieback, the maximum movement of the wall in the 

direction of the cemetery was 0.045 inches in November of 1983. At first glance, a load loss of 9.2 tons appears substantial, 

representing an approximate 25 percent reduction from the lock-off 

load of 40 tons. The question arises as to whether this reduction 

in load was due to creep of the tieback or wall movement. The 

question can be answered by computing the anticipated loss in load 

2-19 
_~__._ _ ---... --.-- 

Figure 17. Time-dependent Redistribution of Strain in the Deformed 

Metal Tube of Tieback 15-8 at 88.1 Tons 

800 

600 

-600 

--- 

t = 0 min. 

t = 700 tilin. 

t = 1000 min. 

t = 1500 min. 

t = 10000 min. 

u 5 10 15 20 25 30 35 40 

DISTANCE FHOM BACK OF- ANCHOR (FT> 
appeared to follow the same pattern observed at Station 304+23. A 

maximum loss in tieback load of 2.4 tons was recorded in the fall 

of 1983. At the elevation of the tieback, the maximum movement of 

the wall in the direction of the cemetery was 0.031 inches. An 

anticipated loss in tieback load of 2.2 tons was computed using 

Equation (2). Since the computed and measured tieback load losses 

are nearly identical, it is concluded that the reduction in 

tieback load was due to wall movement rather than tieback creep. 

A physical interpretation of the data shown in Figure 32 is 

presented below. A drop in water table level behind the retaining 

wall resulted in a reduction of the total pressure acting on the 

wall. A reduction in total pressure resulted in the movement of 

the wall into the hill until a new point of equilibrium between 

the unbalanced forces acting on the retaining was reached. Inward 

movement of the wall was accompanied by an elastic shortening, 

of the unbonded length of the tieback and therefore a reduction i; 

load. 

VII. Conclusions 

Two instrumented, permanent post-grouted tiebacks were installed 

and tested in a stiff-to-very-stiff, red-brown, silty clay at the 

Mt. Carmel Cemetery retaining wall in Baltimore, Maryland, in the 

Winter of 1981-82. Based on the results of the test program, the 

following conclusions are made: 

0 The performance of each tieback satisfied the criteria listed 

in the contract specifications: i.e., 

- the minimum unbonded length was provided, and 

- the creep movement at 75 tons did not exceed 0.1 inches 

between 0.5 and 5 minutes. 

0 Both tiebacks had an ultimate capacity greater than the maximum 

test load of 88.1 tons. 

0 Frictional forces and/or bending stresses were present in the 

tendons of both tiebacks and diminished the magnitude of 

tensile strain measured in the unbonded length at any given 

load. 

0 Distribution of strain, and therefore load, in the anchor 

length of the tendon was non-linear. 

0 Residual strains measured in the anchor of both tiebacks 

increased with increasing load. The opposite behavior was 

observed in the unbonded length and was thought to be due to 

hysteresis of the stressed tendon. 

0 Load transfer is a progressive phenomenon from the front of the 

anchor to the back. A prerequisite for tieback failure may be 

the development of a maximum value of load transfer at the very 

back of the anchor. 

2-21 
6 c---.--..-. _.-_ 1 .---- .-.._____ !.--. ----____ I----.--...L.~~ -____ t -~----.I Q 

3 

- ..---- -._--.-___-- 

~.--I-__----____--._- 

-- 

2-A-15 
0 Piezometers to measure pore water pressures and total pressure 

cells to measure total pressures can be incorporated into a 

test program to investigate the stress field around the 

tieback. 

o More test programs involving permanent tiebacks anchored in 

cohesive soils need to be conducted to evaluate the short-term 

tieback creep tests currently being employed to predict longterm tieback performance. 

o More test programs involving tiebacks anchored in poor soils 

need to be conducted. 

0 Load cells should be used in lieu of or in conjunction with 

strain gauges to measure long-term changes in tieback load. 

2-23 
,. - - 

.; i3 ^. I\ J 

2-A-13 
i- 

\, APPENDIX A 

I 

I--- 

?. 

;----- 

----_ 

2-A-l 
Figure 11. Residual Anchor Movement of Tieback 15-8 

90 

80 

63 

20 

10 

0 

0 0. 05 0. 10 0. IS 0. 20 0. 25 0. 30 

RESIDUAL MOVEMENT <IN. > 

2-A-11 
c 

C 

. . . . I . . . . 

-Nm<lntorAw3 

m 

I\ 

. . w 

. . . . - 

2-A-3 
(9-301x) NIbXLS 

2-A-9 
d-- -p----c 

,,’ ! 

.-t’ ----- 

2-A-5 
Figure 7. Load-displacement Response of a Straight-shafted 

Tieback Anchored in a Cohesive Soil 

90 

10 

0 Q.87.5 tons Q87.5 tons 

0 . s 1 1. s 2 2.5 3 

TOTAL MOVEMENT (IN.) 

2-A-7 
20 

0 

Figure 7. Load-displacement Response of a Straight-shafted 

Tieback Anchored in a Cohesive Soil 

Plastic , 

@ 87.5 tons ' 887.5 tons I 

0 .S 1 1.5 2 2.5 3 

TOTAL MOVEMENT (IN.1 

2-A-7 
e.---- -.- ..-. -. : 1 

I 

i --____ ------A 

,,’ ! 

--- 

.+-I 

2-A-5 
2-A-9 
? 

. . . . . . . . . 

~Nmuln\D~czJ~ 

2-A-3 
90 

80 

63 

20 

10 

Figure 11. Residual Anchor Movement of Tieback 15-8 

0 

0 0.05 0. 10 0. 1s 0. 20 0.25 0.30 

RESIDUAL MOVEMENT <IN.> 

2-A-11 
_.. -~ - ~--. --__ -- - -.-_____._. __._ 

t' igurc 1. Cross-section artd Soil Profile at Station 3041-23 

E. 

I 

. ! 

I 

4 

b. i 

I - 

-1 

: 
\ 

. I 

-- i 

2-A-13 
0 Piezometers to measure pore water pressures and total pressure 

cells to measure total pressures can be incorporated into a 

test program to investigate the stress field around the 

tieback. 

o More test programs involving permanent tiebacks anchored in 

cohesive soils need to be conducted to evaluate the short-term 

tieback creep tests currently being employed to predict longterm tieback performance. 

o More test programs involving tiebacks anchored in poor soils 

need to be conducted. 

0 Load cells should be used in lieu of or in conjunction with 

strain gauges to measure long-term changes in tieback load. 

2-23 
L- .---.- _ ..1- 1 .---I.c-_ - -... i.--. -- I-----_ I ----.- -.L.----. t .--- ----- 

3 

L .----.- --..--” .___-_ -.---l-“--__-_l_ 

----_-.- -- 

2-A-15 
appeared to follow the same pattern observed at Station 304+23. A 

maximum loss in tieback load of 2.4 tons was recorded in the fall 

of 1983. At the elevation of the tieback, the maximum movement of 

the wall in the direction of the cemetery was 0.031 inches. An 

anticipated loss in tieback load of 2.2 tons was computed using 

Equation (2). Since the computed and measured tieback load losses 

are nearly identical, it is concluded that the reduction in 

tieback load was due to wall movement rather than tieback creep. 

A physical interpretation of the data shown in Figure 32 is 

presented below. A drop in water table level behind the retaining 

wall resulted in a reduction of the total pressure acting on the 

wall. A reduction in total pressure resulted in the movement of 

the wall into the hill until a new point of equilibrium between 

the unbalanced forces acting on the retaining was reached. Inward 

movement of the wall was accompanied by an elastic shortening, 

of the unbonded length of the tieback and therefore a reduction iA 

load. 

VII. Conclusions 

Two instrumented, permanent post-grouted tiebacks were installed 

and tested in a stiff-to-very-stiff, red-brown, silty clay at the 

Mt. Carmel Cemetery retaining wall in Baltimore, Maryland, in the 

Winter of 1981-82. Based on the results of the test program, the 

following conclusions are made: 

0 

0 

0 

0 

0 

0 

The performance of each tieback satisfied the criteria listed 

in the contract specifications: i.e., 

- the minimum unbonded length was provided, and 

- the creep movement at 75 tons did not exceed 0.1 inches 

between 0.5 and 5 minutes. 

Both tiebacks had an ultimate capacity greater than the maximum 

test load of 88.1 tons. 

Frictional forces and/or bending stresses were present in the 

tendons of both tiebacks and diminished the magnitude of 

tensile strain measured in the unbonded length at any given 

load. 

Distribution of strain, and therefore load, in the anchor 

length of the tendon was non-linear. 

Residual strains measured in the anchor of both tiebacks 

increased with increasing load. The opposite behavior was 

observed in the unbonded length and was thought to be due to 

hysteresis of the stressed tendon. 

Load transfer is a progressive phenomenon from the front of the 

anchor to the back. A prerequisite for tieback failure may be 

the development of a maximum value of load transfer at the very 

back of the anchor. 

2-21 
1 

(9-301X) NIVtllS 

2-A-17 
After the second logarithmic cycle of the load-hold at 88.1 

tons, the creep rate of tieback 15-8 was double that of tieback 

18-2. The greater creep rate exhibited by tieback 15-8 

appeared to be reflected by the time-dependent changes in 

strain along the deformed metal tube (see Figures 16-18). 

A comparison of the distribution of strain in the tendon of 

each tieback reveals that, for any given load, tieback 15-8 

exhibited slightly higher strains in the unbonded length. This 

behavior is most likely attributable to the presence of greater 

friction forces and bending strains in tieback 18-2. Both 

factors would tend to decrease the magnitude of tensile strain 

measured in the unbonded length. In the anchor length. The 

shape of the strain distribution curves of each tieback is 

different; however, relatively little strain was measured at 

the back of either tendon, even at the maximum test load of 

88.1 tons. The discussion on the mobilization of maximum load 

transfer indicated that a maximum value of load transfer had 

not been developed beyond the mid-point of the anchor length of 

either tendon. Based on the assumption that at failure a 

maximum value of load transfer is developed at the very back of 

the anchor, it appears that both tiebacks had additional 

capacity. A comparison of the distribution of residual strain 

in the deformed metal tube of each tieback reveals that much 

more residual strain was developed at the back of the anchor of 

tieback 15-8. 

B. Long-term Investigation 

The results of the long-term investigation of the two instrumented 

tiebacks are presented together with the long-term performance of 

the retaining wall and soil mass to show that the performance of 

the tiebacks, and the performance of the retaining wall and soil 

mass are interrelated. 

Figure 32 presents data related to the drop in water table level, 

wall movement, and tieback load loss at Station 304+23 and Station 

305+19. At Station 304+23, the data shows that the major drop in 

water table level was accompanied by a significant portion of the 

total wall movement back into the hill in the direction of the 

cemetery. In turn, the largest loss in tieback load occurred 

during the same time period. By November of 1982, the water table 

level has stabilized at an elevation 14 feet lower than its 

original position. After November of 1982, the wall movement 

appeared to be at least partially cyclical in nature, while the 

tieback load loss appeared to have stabilized. A maximum loss in 

tieback load of 9.2 tons was recorded in December of 1983. At the 

elevation of the tieback, the maximum movement of the wall in the 

direction of the cemetery was 0.045 inches in November of 1983. At first glance, a load loss of 9.2 tons appears substantial, 

representing an approximate 25 percent reduction from the lock-off 

load of 40 tons. The question arises as to whether this reduction 

in load was due to creep of the tieback or wall movement. The 

question can be answered by computing the anticipated loss in load 

2-19 
X X 13.9 Tons 

GO0 + 25.4 ” 

v 37.9 ” 

0 50.8 ” 

Ir; * 63.3 ” 

i 0 0 75.4 ” 

-; 400 m m 88.1 ” 

\J 

13.9 Tons 

25.4 ” 

37.9 ” 

50.8 ” 

63.3 ” 

75.4 ” 

88.1 ” 

Figure 19. Distribution of Residual Strain in the Deformed 

Metal Tube of Tieback 15-8 

Figure 19. Distribution of Residual Strain in the Deformed 

Metal Tube of Tieback 15-8 
(2) 

The time-dependent distribution of strain in the 

anchor shows a relatively uniform increase in 

strain at most gauge locations--behavior also 

exhibited in tieback 15-8. This implies that there 

was little time-dependent change in the rate of 

load transfer. The magnitude of strain increase at 

the maximum test load of 88.1 tons, which was held 

constant for 10,000 minutes, was not significantly 

greater than that of any other load-holds which 

were held constant for much shorter periods of 

time. 

Residual-strain distribution in the tendon is 

presented in Figure 27. In the unbonded length, 

residual strain decreased with increasing load, as 

was the case for tieback 15-8. The residual-strain 

distribution in the anchor length conformed to 

expected behavior. As the load increased, the 

residual strain in the tendon increased. The 

transition zone between the unbonded length and the 

anchor length was well-defined and occurred 28 to 

30 feet from the back of the anchor. This 

indicates that the minimum unbonded length was 

provided. 

Deformed Metal Tube 

The distribution of strain in the deformed metal 

tube is presented in Figure 28. Strains tended to 

decrease to zero at the front of the anchor. The 

time-dependent distribution of strain in the 

deformed metal tube shown in Figure 28 exhibited a 

rather uniform increase in strain at each load at 

most gauge locations. In addition, very little 

load was transferred to the back of the anchor, 

even at 88.1 tons. Residual-strain distribution in 

the deformed metal tube is shown in Figure 29. 

Significant compressive residual strains were 

measured at the front of the anchor. The 

compressive residual strains increased as the load 

increased. Very little residual strain was 

measured in the back of the anchor for loads less 

than 88.1 tons. 

C. Computed and Measured Movements 

In the ensuing discussion of computed and measured 

movements, it should be noted that the computed 

movements of the anchorhead would tend to be less than 

the measured movements for reasons identical to those 

presented in the analysis of tieback 15-8 (see Section 

VI.A.1.c). At 88.1 tons, the measured total movement 

and creep movement at the l,OOO-minute mark were used 

because of the shift of the reference point. A 

2-17 
Figure 21. Measured Versus Computed Creep Movement of Tieback 15-8 

90 

80 

70 

60 

G 50 

? 

: 40 -I 

30 

20 

10 

0 

--- Computed 

Measured 

0 0.02 0.04 0.06 0. 08 0. 10 

CREEP MOVEMENT (IN. > 

2-A-21 
2. 

most loads. However, at the higher loads, the measured 

creep movements tended to deviate from the computed creep 

movements. It is highly unlikely that the back of the 

anchor had begun to move through the soil in light of the 

load-displacement-time performance and strain distribution 

characteristics of tieback 15-8. In addition, it should be 

noted that the maximum difference between the computed and 

measured creep movement is on the order of 0.06 inch. 

Tieback 18-2 

a. Load-displacement-time Performance 

Figure 22 presents a plot of load versus displacement 

measured at the anchorhead, while Figure 23 displays 

the residual anchor movement of tieback 18-2. At the 

maximum test load of 88.1 tons, the residual anchor 

movement was only 0.210 inches. Both the elastic loaddisplacement response and the relatively small residual 

anchor movements of tieback 18-2 indicate that it had 

additional capacity beyond 88.1 tons. 

The time-dependent response of tieback 18-2 presented 

in Figures 24 and 25 showed no signs of tieback 

failure. At the maximum test load of 88.1 tons, the 

creep rate was essentially constant at 0.020 inches/ 

logarithmic cycle for the first three logarithmic 

cycles (1 to 1,000 minutes). The creep rate could not 

be determined for the last logarithmic cycle, because 

of the significant decrease in movement throughout this 

cycle. This behavior is more likely attributable to a 

shift in the independent reference point from which 

anchorhead displacement was measured rather than an 

indication of actual physical performance (after the 

l,OOO-minute mark of the load-hold at 88.1 tons, the 

weather suddenly turned warm and caused the ice, which 

partially supported the independent reference point, to 

thaw). The apparent decrease in creep movement also 

masked the true residual anchor movement measured at 

88.1 tons. The true residual anchor movement was most 

likely somewhat greater than the measured 0.210 inches. 

b. Strain Distribution 

(1) Tendon 

The distribution of strain in the tendon of tieback 

18-2 is presented in Figure 26. There was very 

little decrease in strain along the unbonded length 

at most loads. Although bending strains were 

measured in the unbonded length, paired strain 

gauges were used to compensate for them. The 

magnitude of strain measured at the transition from 

the unbonded to the anchor length is interesting in 

2-15 
Figure 23. Residual Anchor Movement of Tieback 18-2 

90 

80 

60 

20 

10 

0 

0 0. ns 0. 10 0. 1s Cl. 20 0.25 0.30 

RESIDUAL MOVEMENT <IN. > 

2-A-23 
The most interesting aspect of the time-dependent 

distribution of strain in the anchor is the 

relatively uniform increase in strain at most 

strain gauge locations for any one load. As a 

result of this kind of behavior, the slope of the 

strain distribution curve between most strain gauge 

locations 

load-hold. 

remained essentially constant during the 

This implies that there was little 

time-dependent change in the rate of load transfer. 

Also note that the magnitude of 

strain increase 

time-dependent 

at the maximum test load of 88.1 

tons, which was held constant for 10,000 minutes, 

was not significantly greater than that of any 

other load-holds which were held constant for much 

shorter periods of time. 

Residual-strain distribution in the tendon is 

presented in Figure 15. In the 

residual 

unbonded length, 

strain decreased with increasing load. 

The reason for this behavior was 

hysteresis 

probably due to 

(see Section IV). The residual-strain 

distribution in the anchor conformed to expected 

behavior. As the load increased, the residual 

strain in the anchor increased. The residualstrain distribution indicates that the minimum 

unbonded length was provided. For all loads, the 

transition zone between the unbonded length and the 

anchor length was well-defined and occurred 27 to 

30 feet from the back of the anchor. 

Deformed Metal Tube 

The distribution of strain in the deformed metal 

tube is presented in Figure 16. The most 

interesting feature of the strain distribution in 

the deformed metal tube is the presence of 

compressive strains at the front of the anchor. 

This also indicates that the load applied at the 

anchorhead was transferred down the tendon to the 

front of the anchor and then transferred to the 

soil; otherwise, compressive strains could not 

exist at the front of the anchor. Also, note that 

the location of the neutral axis (transition 

from compression 

point 

to tension) gradually moved down 

the anchor as the load increased. 

The time-dependent distribution of strain in the 

deformed metal tube shown in Figure 16 differed 

significantly from that in the tendon 

14). During 

(see Figure 

the load-hold at 88.1 tons the 

deformed metal tube exhibited a significant 

increase in strain in the back of the anchor. This 

2-13 
. 

u-2 

N 

(37X,3 Xl-i/ WI> 3iYtl d3383 
Schnabel Foundation Company on instrumented straightshafted tiebacks anchored in 

tieback failure is usually 

clay have shown that 

associated with residual 

anchor movements greater than 0.5 inches. The loaddisplacement response of tieback 15-8 (see Figure 10) 

was much more elastic than that of the straight-shafted 

tieback (see Figure 7). Both the elastic loaddisplacement response and the relatively small residual 

anchor movements of tieback 15-8 indicate that 

additional capacity existed beyond 88.1 tons. 

The time-dependent response of tieback 15-8 ' 

presented in Figures 12 and 13. At the maximum tei: 

load of 88.1 tons, the creep rate was 

constant at 

essentially 

.043 inches/logarithmic cycle in both the 

third logarithmic cycle (100 to 1,000 minutes) and the 

fourth logarithmic cycle 

However, 

(1,000 to 10,000 minutes). 

this creep rate was double the creep rate that 

occurred in the second 

minutes). 

logarithmic cycle (10 to 100 

the third 

The change in creep rate from the second to 

logarithmic cycle has a 

explanation. 

plausible 

load of 

In jacking the load up to the final test 

88.1 tons, the load was inadvertently 

maintained constant for a period of four minutes at a 

load of 86.4 tons. 

load was 

Once the error was discovered, the 

immediately jacked up to 88.1 tons and 

thereafter maintained constant. It is apparent that 

the creep movement of the tieback at 86.4 tons, and the 

effect of the previous loads, influenced the creep 

performance of the tieback at a load of 88.1 tons by 

reducing the amount of creep that occurred in the first 

two logarithmic cycles. The influence of load history 

on the time-dependent performance of tiebacks anchored 

in cohesive soils is been noted in other tests on 

instrumented and non-instrumented tiebacks. 

The creep performance 

imminent failure. 

of tieback 15-8 showed no signs of 

logarithmic 

The maximum creep rate of 0.043 inches/ 

cycle is relatively small for a tieback anchored in cohesive soil. Even at the maximum test load 

of 88.1 tons, the creep behavior was essentially linear 

with respect to the logarithm of time after 

logarithmic cycle. 

the second 

Temperature changes can affect the dial 

gauge support, tendon, and hydraulic fluid in the test 

jack. Temperature-induced movements may become significant 

when anchor creep is small. 

observed during the 

Fluctuations in creep movement 

ton load increment 

fourth logarithmic cycle at the 88.1 

are attributable 

temperature changes between day and night. 

primarily to 

maintain the load constant at 88.1 tons, it 

While trying to 

was 

to adjust the jack pressure 

necessary 

changes. 

in response to temperature 

2-11 
-- - ---...---__ --- , 

2-A-27 
Residual anchor movement is a very useful measure of tieback 

performance, since it is a function of both the tieback 

construction technique and the soil in which the tieback is 

anchored. Reese (1976) highlighted the fact that the 

displacements required to mobilize ultimate capacity in bearing 

are much larger than the displacements required to mobilize 

ultimate capacity in shear in most soils. A large number of tests 

conducted on tiebacks using extensometers have verified that the 

anchor moves when loaded and partially rebounds when unloaded, and 

that the residual anchor movement can be measured reliably in a 

cyclic test such as the performance test (Schnabel, 1982). Since 

larger anchor displacements are required to mobilize greater 

tieback capacity (for loads less than ultimate), residual anchor 

movement tends to increase as the applied load increases. To 

date, no acceptance criteria exists to evaluate tieback 

performance based on residual anchor movement. 

Time-dependent tieback displacements at constant load are used to 

evaluate the long-term performance of a tieback. The timedependent or creep movements typically are plotted on a semilogarithmic scale as a function of time (see Figure 8). For each 

load, the creep movement per logarithmic cycle (or creep rate) is 

determined by computing the slope of a best fit straight line 

drawn through the data points. If a straight line cannot be 

drawn, the creep rate is determined by computing the maximum slope 

formed by the data points. Creep rate can then be plotted as a 

function of load. To date, tieback specialists cannot agree on 

the acceptance criteria to be used to evaluate tieback performance 

based on time-dependent movements. 

The preceding discussion has focused its attention on the 

measurements used by the engineering profession to evaluate a 

tieback's performance. Elastic movements, residual anchor 

movements, and time-dependent movements are used to assess both 

the short-term and long-term load carrying capability of a 

tieback. The ensuing discussion presents a step-by-step 

description of the analysis used herein to evaluate a tieback's 

performance based on computed performance which can provide 

insight into the measured performance. 

The most obvious contribution of the computed performance of each 

tieback is the determination of the distribution of strain, and 

therefore loadl, along the tieback for a given applied load and a 

given time (see Figure 9). The shaded area represents the total 

amount of strain that developed during a particular load-hold. 

The strain distribution in the unbonded length can verify easily 

the provision of a minimum desired unbonded length--a criteria 

used by the entire engineering profession to evaluate a tieback's 

performance. The installation of strain gauges at critical 

locations along the unbonded length, e.g., outside and inside the 

face of the tied back structure and at the transition from the 

unbonded length to the anchor length, can provide valuable 

information concerning the influence of friction and bending on 

the distribution of strain in the unbonded length. Typically the 

2-9 
400 

300 

200 

100 

0 

-100 

-200 

-300 

-400 

Figure 29. Distribution of Residual Strain in the Deformed Metal Tube of 

Tieback 18-Z. 

* 13.9 Tons 

* 25.4 ” 

l 37.9 ” 

v 50.8 ” 

A -1 

/r/ 4, 

0 ” 

0 63.3 754” 

/ 

/ 

/ 

/ 

0 5 10 15 20 25 30 35 40 

DISTANCE FROM BACK OF ANCHOR (FT.> 
In setting up each tieback test, extreme care was exercised to 

ensure that the bearing plates, load cell, and jack were aligned 

properly. This was done in order to minimize the effects of 

friction and eccentric loading due to mis-alignment. Prior to the 

start of each test, each strain gauge was calibrated internally 

and initialized to give a "zero" reading to which all subsequent 

strain gauge readings were referenced.' A 5.5-ton alignment load 

was then placed on the tieback to seat and align the testing 

hardware. At this time, the dial gauges were set up to measure 

movement of the anchorhead. All subsequent dial gauge 

measurements were referenced to the initial "zero" reading at 5.5 

tons. The test proceeded by setting the desired load on the 

Vishay P-350A strain indicator. The jack was then pumped up to 

the desired load. At this time, the dial gauge and all strain 

gauges were read. If the next load was to be held constant, the 

jack was pumped up to the desired load within one minute. An 

initial reading was taken when the desired load was first reached. 

The one-minute reading of the load-hold corresponded to the time 

that had elapsed since the pump was first activated. That is, the 

time lapse between the initial '1 zero '1 reading and one-minute 

reading of the load-hold was always less than 60 seconds. During 

the load-hold, it was necessary to adjust the hydraulic pressure 

in the test jack in order to maintain the load constant. Load was 

decreased to the alignment load after each load-hold by setting 

the Vishay P-350A strain indicator to the alignment load and 

opening up the release valve on the pump. At the end of each 

test, the load was reduced to zero and all strain gauges were read 

and checked for internal calibration. Subsequently, each tieback 

was locked-off at 40 tons by reloading the tieback to 40 tons, 

tightening a nut against the bearing plate, and releasing the load 

from the jack. 

B. Two-year Monitoring Program 

The two-year monitoring program consisted of monitoring the change 

in load in each of the tiebacks at least once every three months 

during this period of time. The strain gauge readings taken 

immediately after each of the tiebacks was locked-off at 40 tons 

served as the initial zero to which all subsequent readings were 

compared. 

The procedure used to take a set of readings is described below. 

The dummy gauge was used to zero each of the 30 channels on the 

Vishay 220 readout unit prior to visiting the site. At the site, 

the dummy gauge was used again to re-zero the 30 channels. 

Subsequently, the strain gauges on both instrumented tiebacks were 

read within a period of 60 minutes. 

Note that the strain gauge readings taken in June of 1982 are 

omitted from this report because they were considered unreliable. 

The reason for their un-reliability was traced to the failure to 

pot the military connectors which served as the junction between 

the strain gauges and the Vishay 220 readout unit. Grouting of 

2-7 
Figure 31. Measured Verjus Computed Creep Movement of Tieback 18-2 

33 

20 

10 

0 

-w- Computed 

Measured 

0 0.02 0.04 0.06 0. 08 0. 10 

CREEP MOVEMENT (IN.> 

2-A-31 
0 a calibrated 150-ton capacity load cell to monitor load at the 

anchorhead during the short-term investigation; 

0 a calibrated test pump and jack to load the tieback; 

0 two dial gauges to monitor movement of the anchorhead relative 

to a fixed point to the nearest .OOl inch; 

0 a Vishay 220 digital strain readout unit capable of reading and 

recording the response of 30 strain gauges continuously at a 

rate of one strain gauge per second; 

0 a Vishay P-350A strain indicator to monitor load in the load 

cell; and 

0 a dummy gauge used to zero the Vishay 220 digital strain 

readout unit prior to each set of readings over 

monitoring period. 

the two-year 

In general, the strain gauges 

initial two-week test program. 

performed quite well during the 

less 

Strain gauge failure occurred in 

than 17 percent of the strain gauges. However, the 

performance of the same strain gauges over the two-year monitoring 

period was less satisfactory. This type of performance is to be 

expected from electrical resistance strain gauges which are 

accurate but, 

very 

due to electrical aging of components and creep in 

bonding agents, tend to "drift" with time. 

a hostile 

Prolonged exposure to 

environment over a period of time was probably the 

reason for the majority of the strain gauge failures. 

Since the original intention of the two-year 

was to monitor changes in 

monitoring program 

tieback load, the strain gauges of 

primary interest were those located in the unbonded length of each 

tieback. Fortunately, 8 out of the 10 strain gauges located in 

the unbonded length functioned properly for the entire 

monitoring period. Therefore, 

two-year 

the two-year load change data 

presented in Section VI is considered reliable and accurate 

provided that certain limitations associated with the use of 

strain gauges are recognized. First, electrical 

resistance strain gauges to 

the tendency of 81 drift II II with time influences the 

absolute magnitude of load measured more so than the trend in load 

change. Both the strain gauges installed on the two tie backs and 

the dummy gauge used to zero the Vishay 220 unit were affected by 

this "drift" phenomenon. Second, 

of elasticity of the 

the variation in Young's modulus 

convert strain to load 

introduces 

bar required to 

error to the absolute magnitude of load measured. The conversion of strain to load required the use of Equation (1): 

P = eEA . ..(l) 

where P = load 

e = strain (x 10 -6 ) 

E = Young's modulus of elasticity 

A = cross-sectional area of the tendon 

2-5 
REPORT NO. 3 

PERMANENT TIEBACK ANCHORS IN COHESIVE SOIL 

DEMONSTRATION SITE 

SR-90 (SR-5 to CORWIN PLACE) 

SEATTLE, WASHINGTON 

bY 

Hart Crowser and Associates, Inc. 

WASHINGTON STATE 

DEPARTMENT OF TRANSPORTATION 

MARCH 1984 
B. Installation 

The installation procedure consisted of four major operations (see 

Figure 4): 

o punch an 8-l/2-inch diameter hole through the wall; 

o drill hole through soil to desired depth and insert TMD 

assembly; 

o grout the deformed metal tube to the soil; and 

o grout tendon to the deformed metal tube. 

An 8-l/2 inch diameter hole was punched through the wall by using 

an 8-inch diameter down-the-hole hammer with a button bit attached 

to its end. Upon encountering the first section of rebar 2-l/2 

inches from the front face of the wall, a laborer used a torch to 

burn off the rebar. The second section of rebar, located near the 

back face of the wall, was burned out with a lance. 

Once the hole had been punched through the wall, the hole was 

drilled to a final depth of 57 feet using a 6-inch diameter auger 

with a clay bit attached to its end. From the experience gained 

during the installation of the 96 non-instrumented tiebacks, it 

had been decided the hole could be drilled uncased. This greatly 

simplified the drilling operation. Very soft, soupy soil with 

running sand was encountered while drilling the hole for tieback 

18-2. Fortunately, this presented no problems during the 

installation of the tieback. 

Once the auger had been extracted from the drill hole, the 

deformed metal tube, with grout valves located every 3.3 feet 

along its length, and the PVC pipe were placed in the hole. The 

assembly was centered in the hole by means of plastic centralizers 

attached to the deformed metal tube. After the TMD assembly had 

been inserted in the drill hole, a 3-inch diameter double packer 

was positioned inside the deformed metal tube opposite the bottom 

grout valve. The packer was inflated and grout subsequently was 

pumped through the last valve until the grout could be seen 

exiting from the drill hole. The packer was removed and the 

inside of the TMD assembly was flushed with water and blown out 

with compressed air until the effluent was fairly clear. The 

grout was then allowed to set up overnight. 

The post-grouting operation began on the next day. The packer was 

positioned inside the deformed metal tube opposite the bottom 

grout valve and inflated. Grout was pumped slowly through the 

valve until either a maximum grout pressure of 40 bars was reached 

or one bag of cement had been pumped through the valve. The 

quantity of grout pumped through each valve was limited, because 

2-3 
TABLE OF CONTENTS 

I. INTRODUCTION ............................................ 3-l 

II. SUMMARY ................................................. 3-2 

III. SUBSURFACE EXPLORATIONS AND LABORATORY TESTING .......... 3-2 

IV. 

V. 

VI. 

TIEBACK INSTALLATION .................................... 3-3 

PHASE I TESTING: PULL-OUT AND SHORT TERM CREEP ......... 3-4 

A. Ultimate Anchor Testing .............................. 3-4 

B. Creep Testing ........................................ 3-4 

PHASE II TESTING: LOCK-OFF MONITORING AND LONG TERM .... 3-6 

MONITORING 

A. Lock-Off Loads ....................................... 

B. Creep Criteria ....................................... :I,: 

C. Load/Deflection Criteria ............................. 

D. Long Term Monitoring ................................. 

VII. CONCLUSIONS . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3-10 

3-i 

PERMANENT TIEBACK ANCHOR DEMONSTRATION PROGRAM 

SR-90 (SR-5 TO CORWIN PLACE) 

I. INTRODUCTION 

This report provides the results of the first 100 days of data of 

the SR-90 Permanent Tieback Anchor Demonstration Program. The 

purpose of the study was to establish creep characteristics of 

non-pressure grouted tieback anchors installed in the Seattle 

over-consolidated silts and clays in a predesign testing program. 

The work included on-site explorations, laboratory testing, 

installation of nine tieback anchors, and load-creep testing of 

the anchors. Our work has been accomplished in accordance with 

our final design submittal authorized in Agreement Y-2573 

Supplemental dated June 24, 1983 (FHWA Work Order DTFH 71-83-931- 

WA-04). 

Selection of the tieback anchor for this study was based on the 

standard of practice throughout the Pacific Northwest. With only 

minor exceptions, the tieback anchors used throughout the area are 

non-pressure grouted anchors which typically ranged in diameter 

from 12 to 18 inches. For this study, 12-inch nominal diameter, 

20-foot long, non-pressure grouted anchors were utilized. 

The site of the study is located at the western end of the SR-90 

project, near the SR-90, SR-5 interchange, as shown on Figure 1. 

The site was selected to utilize one of the existing cantilevered 

cylinder pile walls as a reaction to the tieback anchor loads. 

The lo-foot diameter reinforced concrete cylinder piles provided a 

very rigid reaction, and for all practical purposes eliminated the 

influence of backfill loads from the testing program. 

The tieback testing for this study was completed in two phases. 

In the first phase, two anchors were loaded incrementally to pullout. Each load level was held for up to an hour, while monitoring 

bar displacement to establish creep versus load characteristics. 

Based on that data, anchor loads were selected for long term creep 

monitoring. During the second phase of the test, seven anchors 

were loaded to and locked off at various percentages of the 

ultimate anchor capacity (20 to 60 percent). The anchors are 

currently being monitored to establish long term creep versus time 

relationships as a function of anchor load. 

The body of the report provides a discussion of the various 

aspects of the project, including subsurface exploration and 

laboratory testing, tieback installation, ultimate anchor testing, 

and creep data interpretation. 

3-l 
Z-E 

'III 33ViMlSEI~S SNOIiLV?JO'IdX3 CIN'V AtlOtT?JlIOEWI 9NI&S3& 

8'0 X SL = 09 sd?x l 

0 aq& umur~xtm 6ur5p0M uoTsua7 (Mna) SAM pa7eTnDTeci 7e 8.0 ,DJ ' x0 
fracture zones 

soil as the 

project. 

0 The Plasticity Index (PI) of the soil ranged from 17 to 39 

percent, with a majority of the samples in the range of 24 to 

were noted. We have typically classified this 

Older Marine unit in our reports for the SR-90 

32 percent. The liquidity index ranged from 0.09 to 0.61. 

Soil with a PI greater than 20 percent generally requires an 

assessment of creep potential. Moisture contents of samples of 

drill cuttings from the anchor zones of the tiebacks, taken 

during tieback drilling, were in the range of 33 to 39 percent. 

0 The unconsolidated undrained triaxial tests indicated shear 

strengths in the range of 1600 to 5300 psf. Strain at failure 

was generally in the range of 2 to 4 percent. The results of 

strength testing within the Older Marine soils conducted for 

this and other projects would indicate that a definite trend of 

increased strength with depth does not necessarily exist. The 

strength of these glacially consolidated materials will tend to 

vary from point to point within the soil mass. 

IV. TIEBACK INSTALLATION 

The anchors were installed on September 20 and 21, 1983, The 

locations of the anchors are shown on the Site and Exploration 

Plan, Figure 2. Generalized subsurface profiles at the site are 

shown on Figures 3 and 4. These figures also show the 

configuration of the test program in, relation to the subsurface 

conditions disclosed by the explorations. Figure 5, the Typical 

Tieback Installation, indicates the configuration and dimensions 

of the tiebacks as installed for this project. The following is a 

summary of the tieback installation program. 

The anchors were installed usina a 12-inch nominal diameter LDH 

auger rig. The actual drilled hole diameter ranged up to 14 

inches, which is typical for the method of installation. The 

average overall length of each tieback was 55 feet. The 

grouted anchor length was approximately 20 feet, 

The holes were drilled at an inclination of 25 to 30 degrees 

from horizontal. Flatter angles of 20 to 25 degrees were not 

generally achievable due to site geometry limitations (site 

width, position of concrete footing). 

The tieback bars consisted of 1 3/8-inch diameter Dywidag bars. 

The bars were installed and anchors grouted by open-hole 

methods. The anchors were not pressure-grouted. 

The anchors were installed in the Older Marine unit (hard silts 

and clays). The holes were generally dry with occasional 

localized saturated zones. Tieback TB-5 encountered a 

saturated zone, apparently within the unbonded length. Flowing 

water exited this hole shortly after drilling, and continued to 

3-3 

and 13). The slope of each curve, designated the creep 

coefficient (UC), was then plotted versus the corresponding load 

level. The typical shape of the creep coefficient curve in this 

study is nearly horizontal fo'r loads lower than approximately 50 

kips, at values of usually less than 0.03 inches per log cycle of 

time. At higher loads the value of increases with each load 

level, with the creep coefficient curve sweeping upward (See 

Figures 7, 8, and 9). 

The critical creep tension for each anchor was determined from the 

creep coefficient plot. Two methods of calculating the critical 

creep tension were utilized, as follows: 

Tc: The load level where the slope of the creep coefficient curve 

becomes tangent to the initial straight line portion of the curve. 

Tc': The load level defined by the intersection of the straight 

line extended from the initial portion of the curve, with the 

extension of the straight line extended from the latter portion of 

the curve. 

Log Tim 

8 i 5 g 0 t 8 8 l&l-L TC Tti 0 01 0.. 0.6 0.2 

Anchor Load 

To 

A. CREEP CURVES B. CREEP COEFFICIENT PLOT 

(Note: Figure 7 also presented following text) 

The following is a summary of the sequence of creep testing for 

the Phase I anchor testing. 

0 Tieback U-l: Short term creep data was collected while testing 

ultimate anchor U-l to pull-out. The creep curves and critical 

creep tension curve are presented on Figure 8. The data 

indicates a critical creep tension of : Tc' = 79 kips, Tc=80 

kips. 

0 Tieback U-2: The creep data for the test is presented on Figure 

9. The creep coefficient showed a significant rise from the 

first 70 kip load to the 80 kip load, increasing from 0.02 to 

0.06. The data indicated critical creep tension of: Tc'=75 

kips, Tc=70 kips. Following the 80 kips short term hold, the 

load was decreased to 70 kips for the long term hold. The load 

was maintained for 32 hours, at which time an equipment 

malfunction resulted in the load increasing to 105 kips. The 

creep data obtained during the first 32 hours did not plot as a 

straight line, but rather had an upward curvature on the 

displacement vs. log-time curve. The curvature may be due to 

3-5 
9-E 

3NItIO;GINOR 

'IA 3SWd II :3NI&SBt 660-)I301 3NI?lO&INO~ CINV SNOT FRI3lL 
the 0.8 Tc' criteria. The creep coefficients (d) were less then 

0.08 inches per log cycle of time for all of the loads up to 80 

kips. With the exception of the second 70 kip load on Tieback U2, the creep curves plotted as straight lines on the semilog 

graphs. 

c. Load/Deflection Criteria 

Load/Deflection criteria for anchor tests is based on elastic 

elongation of the tendon or bar during stressing. The criteria 

for minimum and maximum deflection is described in the draft of 

the FHWA manual for Demonstration Project Number 68, Permanent 

Ground Anchors. The minimum deflection of the anchor head is 80 

percent of the theoretical elastic elongation of the free, 

unbonded tendon or bar length. The maximum deflection at the 

anchor head is the theoretical elastic elongation of the tendon or 

bar length from the jack to the center of the bond length. The 

center of gravity of the bond stress should not be beyond the 

midpoint of the bond length, if the maximum deflection criteria is 

met. 

The anchor test data was evaluated for these criteria based on a 

steel area of 1.485 square inches for the bar and a modulus of 

elasticity, E, of the tendons, 29 x lo6 p.s.i, A free, unbonded 

bar length of 32 feet, and a length of 42 feet to the center of 

the bond length were used. The anchor tests all met the minimum 

deflection criteria. Five of the tests exceeded the maximum 

deflection criteria by the following percentages. Of the 60 and 

70 kip tests, two of five exceeded the criteria, by 10 and 12 

percent of the theoretical elongation. Of the 42 kip tests, one 

of two exceeded the criteria, by 28 percent. Of the 24 to 28 kip 

tests, two of three exceeded the criteria, by 37 and 65 percent. 

A possible explanation of the deflections observed which exceeded 

the maximum criteria could be inaccuracy of relative deflection 

measurements due to straightening of the system with increasing 

loads. Another could be inaccuracy (apparently at low loads) of 

load measurement. Because the percentage of deflection in excess 

of acceptance criteria is higher for the lower loads, it appears 

that the center of gravity of bond stress is not beyond the 

midpoint, as this would be unlikely for the 24 kip tests. 

Further analysis of this phenomena will be made during review of 

load transfer mechanisms as indicated by additional strain gages 

installed within the anchor zone of selected anchors. This 

analysis will be completed and forwarded under separate cover. 

The selected value of Tuw for determining lock-off loads was 60 

kips. The value was selected following review of the data and 

discussion with representatives of both FBWA and WSDOT. The 

actual tieback lock-off loads are presented on the following 

table. 

3-7 
8-E 

ZP 

SOT 

ZP 

L.-t1 

LP 

OL 

EL 

EO'C 

E”: 

sz 

OL 

82 

ZP 

VP 

z9 

L-8lTl 

9-0J.l 

S-ELTI 

p-m (-pU?d 

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The load was monitored with the hydraulic load cell at the waler, 

and with a vibrating wire strain gage spot welded to the Dywidag 

bar. The vibrating wire strain gage was calibrated on the l-3/8 

inch diameter Dywidag bar. 

x lo-6), 

A strain change of 20 microstrain (20 represents a load change in the bar of 1.0 kips. The 

deflections ( & ) in the no load zone were estimated from the 

vibrating wire strain gage data as follows: 

where L is the strain recorded from the gage, and L is the 32 

foot length of the no load zone. Based on this relationship, a 

deflection (A) of 0.01 inches would be represented by a 

recorded strain ( e ) of 26 microstrains. That level of strain 

would also represent a 1.3 kip change in the load of the anchor, 

based on the above calibration. 

Following lock-off, the readings were obtained on a one to two 

week basis, with the exception of the first week, where data was 

collected on the first, second, and fourth day. The monitoring to 

date has been based primarily on the vibrating wire strain gage 

data. 

Based on published information of time-dependent relaxation of 

prestressing steel presented in FHWA Report Number FHWA/RD-82/047, 

and on verbal information from Dywidag System International, we 

have not subtracted an allowance for load relaxation (or "steel 

creep") from our long-term creep monitoring data. The published 

information, applicable only to bars stressed to 70 percent of 

ultimate strength (140 kips for this project), indicates load 

relaxation would be on the order of 4 percent for the loo-day 

duration of the monitoring to date. Dywidag representatives 

inform us, however, that at loads lower than 70 percent of 

ultimate (bars for this project are stressed at 10 to 30 percent 

of ultimate), load relaxation would be minimal and probably less 

than 1 percent. In the absence of published data applicable to 

this project, we have not considered load relaxation in our data 

due to its anticipated low magnitudes. 

The load dissipation observed during long term monitoring is due 

to creep of the anchor. Load dissipation to a small extent occurs 

in the anchors as the anchor creeps under the applied load, and 

the strain in the bar is reduced. The creep coefficients measured 

during long term monitoring therefore are affected by the 

reduction in load that occurs as a result of anchor creep, and are 

lower than the short term creep coefficients. 

To date, load dissipation has been relatively small, and little if 

any I affect on creep coefficient has been noted. It is not known 

at what level of load dissipation the coefficients would be 

significantly affected. For purposes of this program the tiebacks 

should be restressed up to the initially applied lock-off loads if 

the tieback load dissipates by more than 10 percent of the lockoff value. If load dissipation on the order of 10 percent is 

3-9 
01-E 

'S;CO~XIB lQTM B a6uez 30 ~00'0 07 zTo.0 Saq=>UT zad 

- _-- 
log cycle of time for the 62 to 63 kip anchors. The 70-kip 

anchor has a creep coefficient of 0.009 after being locked off 

about 20 days. 

0 The creep curves tended to show an initial period of virtually 

no movement, followed by a downward sloping curve. The break 

in the curve could be due to the preloading effect from the 

short term creep test, or other yet undescribed factors. 

0 The coefficients for load drop off tended to range between 0.1 

to 0.5 kip per log cycle of time for a 25 kip anchor, and up to 

1.5 kips per log cycle of time for the 63 and 70 kip anchors. 

None of the loads have been re-applied since locking off the 

anchors. 

0 The data obtained to date for the anchors exhibits some 

scatter, and continued monitoring is necessary to evaluate the 

long-term anchor creep coefficients. We recommend that 

monitoring continue at least until October, 1985 which would 

provide 2 years of monitoring since lock-off. As shown on 

Figure 15, the additional data would include approximately one 

log cycle of time from the 100 day mark. 

0 The data collected to date indicates that a 60 kip load on the 

demonstration anchors satisfied the maximum working tension 

criteria. 

0 According to the recent lock-off data from the 70 kip load on 

TB-4, 70 kips may also satisfy the maximum working tension 

criteria. However previous data from the ultimate test on 

Tieback U-2 indicates that 70 kips may not satisfy the maximum 

working tension criteria. With lack of additional data, our 

recommendation would be to continue to use 60 kips as the 

maximum working tension for the project. 

3-11 


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IJI Split Spoon Sample Location and Number 

t.il Shelby Tube Sample Location and Number 

P Sampler pushed hydraulically, not driven. J-712 -12 March 1984 

HART-CROWSER %I associates inc. 

Figure.3 

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Typical Tieback Installation 

Extensometef Head Detail: 

A) Phase I, Testing 

SI&II~SS steel Exteneometer Head 

#I (Red) Bolted to Fleterence Hesd 

+2 (White), and #3 (6lue) free to slide 

B) phase II, Long Term Monitoring 

+ VI~IW 

~~*fRcD) 

+ Only installed on TB-6, U-l and U-2 

PVC Sleeve 

Stainless Steel Extensometer Head 

#I (Red) free to slide 

x Bolted to Reterence Head - Anchored 

behind Waler to the ground 

#2 (White). and #3 (Blue) tree to slide 

J-712-12 March 1984 

HART-CROWSER & associates inC. 

Figure 5 
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1 > to*1 min to*1 0 ta6Omin 

Log Time 

A. Creep Curves 

0 0.2 0.4 0.6 0.8 

Anchor Load 

To 

B. Creep Coefficient Plot 

T,= Anchor Pull Out Load 

Critical Creep Tension: 

Tc= The load level where the slope of the creep coefficient curve 

become8 tangent to the initial straight line Portion of the curve. 

Tc’= The load level defined by the intersection of the straight line 

extended from the initial portion of the curve, with the extension of 

the Etraight line extended from the latter portion of the curve. 

J-712-12 March 1984 

HART-CROWSER & associates inc. 

Figure 7 

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Tieback U-2 

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2.0 

ii c 

2 

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4 

K 

5 

5 ‘L 

1 

i 1.0 

3 

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50 Kips, 

40 Kips 

30 Kips 

20 Kips 

10 100 

Time in Minutes 

1000 

LOAD 

IN KIPS 

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110 

105 

100 

90 

80 

80 

70 

70 

60 

10000 

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Tc’+ 75 kips 

al 

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CYCLE OF TIME 

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.02 

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kips 

0 Short-term Hold 

Q Long-term HoJd 

1, Short-term Hold 

on Reload 

Following’initial 

106 Kip Load 

40 60 

Load in Kips 

3-A-9 

80 100 110 

J-712-12 March 1984 

HART-CROWSER & associates inc. 

Figure 9 
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Tieback TB-3 

42 Kips 

1.0 

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E 

Q 0.5 

K 

ii 

Tieback TB-5 

0 

24 Kips 24 .Ol 

10 Kios 10 0 

lJ 

LOAD 

IN KIPS 

42 

30 

20 

10 

1 10 100 

Time in Minutes 

1000 10000 

at 

INCHES/LOG 

CYCLE OF TIME 

.02 

.Ol 

.Ol 

.Ol 

Short-term Hold 

l TB-3 

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Load in Kips 

J-712-12 March 

3-A-11 

HART-CROWSER & associates Inc. 

Figure 11 





REPORT NO. 4 

PERFORMANCE MONITORING OF A HIGHWAY TIEBACK WALL 

DEMONSTRATION SITE 

I-71, CARROLL COUNTY, KY-227 

CORRECTIVE LANDSLIDE MEASURE 

bY 

Bobby W. Meade 

Research Investigator 

David L. Allen 

Chief Research Engineer 

and 

Tommy C. Hopkins 

Chief Research Engineer 

KENTUCKY TRANSPORTATION RESEARCH PROGRAM 

COLLEGE OF ENGINEERING 

UNIVERSITY OF KENTUCKY 

LEXINGTON, KENTUCKY 

JULY 1986 
The contents of this report reflect the views of the authors who 

are responsible for the facts and the accuracy of the data 

presented herein. The contents do not necessarily reflect the 

official views or policies of the Kentucky Transportation Cabinet, 

the Federal Highway Administration, nor the University of 

Kentucky. This report does not constitute a standard, 

specification, or regulation. 
TABLE OF CONTENTS 

I. PROBLEM STATEMENT . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3-3 

A. Geology . . . . . . . . . . . . . . . . . . . . . . . ..*.................... 

B. Subsurface Conditions Prior to Corrective Action . . . . . 

II. REMEDIAL OPTIONS . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3-4 

III. STUDY PROPOSAL . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3-4 

IV. WALL DESIGN, LAYOUT, AND CONSTRUCTION ................... 3-5 

A. General Construction ................................. 3-5 

B. Tiebacks ............................................. 3-6 

C. Instrumentation ...................................... 3-7 

V. TIEBACK TESTING . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3-9 

VI. TEST DATA AND RESULTS .................................. 3-10 

A. Laboratory Data ..................................... 3-10 

B. Lateral Movements ................................... 3-10 

VII. PRESSURE DATA .......................................... 3-11 

A. Water Table ......................................... 3-11 

B. Durability .......................................... 3-12 

VIII. PROBLEMS . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3-12 

IX. CONCLUSIONS . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3-12 

4-i 

I. PROBLEM STATEMENT 

Over the past several years', an unstable cut-and-fill embankment 

on KY 227 (Carrollton-Worthville Road) in Carroll County had been 

failing. The problem area .is located approximately eight miles 

southeast of Carrollton between Stations 234+25 and 244+25. This 

slide caused numerous maintenance problems for the roadway, which 

was frequently overlaid, a:?d for the nearby railroad track, which 

had to be periodically realigned. As shown in Figure 1, 

slippage had lowered the shoulder several feet. 

prior 

A view from 

approximately the same location is shown in Figure 1A. 

of the slide area, is shown in Figure 2. 

A vicinity 

map 

vicinity map of the slide area. 

Figure 2 presents a 

A. Geology 

This site lies in the 

topographic region of 

interbedded shales and 

of Ordovician age. 

present in the river flood 

part of the Outer Bluegrass 

It consists of predominately 

of the middle and upper series 

and recent alluvium are 

The primary geologic involved in the slide is 

Formation. 

the Kope 

shale interbedded 

Kope Formation is a medium gray 

gray limestone. Limestone 

generally comprises about' 15 to 30 percent of the formation and 

usually occurs in even to sbightly irregular beds about 12 inches 

thick. Thin beds of laminated calcareous siltstones are also 

occasionally found. ' 

B. Subsurface Conditions P/tzior to Corrective Action 

The Division of Materials ok the Kentucky Department of 

conducted a geotechnical ~ exploration of 

Highways 

located at 

the site with borings 

Station 236+00 

Station 236+00 

--~ 45 feet right of centerline, 

Station 236+50 

--I 45 feet left of centerline, 

--i 50 feet right of centerline, and 

Station 237+00 --I 45 feet left of centerline. 

Laboratory tests were condu ted on obtained at Station 

236+50 -- 50 feet right of 

samples 

enterline. Test results indicated the 

material to be an A-7-6 ~ (19) soil according to the AASHTO 

classification system a@ classification 

a CL according to the Unified 

system. The natural moisture content was 17 

percent at 

21.5 feet, the AASHTO 

to 11.5 feet. At a depth of 20.0 to 

classification 

was A-6 (18) and the Unified 

was CL. natural moisture content was 16 

percent. 

Slope inclinometers number 5 and 2 were installed at Station 

236+50 (50 feet right o 

feet right of centerline), 

centerline) and at Station 236+00 (45 

espectively. Both slope inclinometers 

4-l 
indicated deflection rates of 0.2 inch per month. The sliding 

plane was at elevation 466.0 feet, which was 25 feet below the 

shoulder. 

The two borings left of centerline were used as observation wells. 

The average water-table depth was 2 feet at Station 236+00 (well 

W and 4 feet at Station 237+00 (well 1B). A plan view of the 

site, including structures, natural features, and instrumentation 

previously discussed, is shown in Figure 3. 

The in situ soil strength parameters were estimated by a back 

analysis iteration. The soil unit weight and cohesion were held 

constant while the angle of internal friction was varied to arrive 

at a safety factor of 1.0. The results are 

W= 128 pcf, 

C = 0 psf, 

0 = 18 degrees. 

The cross section used in the existing conditions analysis is 

shown in Figure 4. 

II. REMEDIAL OPTIONS 

Several remedial procedures were considered, with an increase of 

at least 30 percent in the existing safety factor as the critical 

criterion. Realignment of KY 227 further into the hillside was 

readily eliminated as impractical. That option would involve 

disturbing an already marginally stable hillside and require 

excavation, re-paving, and right-of-way changes for at least onehalf mile. Flattening the slope also was eliminated because of 

adverse effects on the railroad. 

Three other remedial procedures were considered; horizontal 

drains, rail piles, and a tied-back control wall. The horizontal 

drain option would result in a safety factor of 1.1 to 1.2 at an 

approximate cost of $62,000. This did not meet the criterion of a 

30-percent increase of the safety factor. Rail piles perform best 

at depths to bedded material of less than 15 to 20 feet, The soil 

depths at this site exceed 20 feet. The tied-back control wall 

was chosen as the best alternative. A contract for construction 

of the wall was awarded on December 18, 1983. 

III. STUDY PROPOSAL 

The use of tieback walls to control landslide problems is somewhat 

limited. For that reason, a study was initiated with the 

objectives of: 

(1) documenting construction procedures and obtaining shortterm and long-term experimental data on tieback wall 

performance, 

(2) analyzing field behavior using instrumentation installed 

on and near the wall, and 

4-2 
(3) making recommendations as to the effectiveness and 

future use of tieback walls constructed with treated 

wood lagging for correcting highway embankment sliding. 

Monitoring of the wall is to continue for a period of five years. 

Data and observations subsequent to this report will be presented 

annually in the form of a memorandum. 

IV. WALL DESIGN, LAYOUT., AND CONSTRUCTION 

The tieback wall was designed and constructed by the Schnabel 

Foundation Company in compliance with Department of Highways 

Special Notes. Design personnel were supplied a design restraint 

force of 33,000 pounds per linear foot of wall where the sliding 

plane was 25 feet below the top of the wall. A safety factor of 

1.5 was assumed. This translates to a normalized uniform loading 

(P) oft 

P = 0.0528 kip/foot x h, 

where h = the height of the wall in feet. 

The wall was constructed using a system of steel H-piles, 

pressure-treated wood lagging, and corrosion-protected tiebacks. 

A typical wall section is shown in Figure 6. The assumption of a 

safety factor of 1.5 and testing up to 133 percent of design load 

(to be discussed in the TIEBACK TESTING section) resulted in test 

loads of 200 percent of expected loading. This conservative 

approach was probably the result of a lack of experience with this 

type structure. 

The tied-back wall, as designed, extended from Station 234+25 to 

Station 244+25. At Station 234+25, the wall was 76 feet right of 

centerline. The wall gradually approaches centerline, and at 

Station 235+50 it is 70 feet right. At that point, the wall bends 

toward the centerline, making it only 40 feet right at Station 

236+00. From Station 236+00 to the ending station, the wall is 

approximately 43 feet right of the centerline. The height of the 

wall is approximately 15 feet above finished grade. Total cost of 

the wall and associated efforts was $483,000. A plan view and 

front view of the wall are shown in Figure 7. Excluding the cost 

of excavation, cost of the wall was $31.25 per square foot of 

exposed wall. 

A. General Construction 

A total of 126 soldier piles were driven on approximately 8-foot 

centers. Of these, 90 were 10 x 42 H-piles and the remaining 36 

were 12 x 53 H-piles. Driving points were specified for 

penetration of a boulder zone and to insure proper seating. Piles 

were driven to a resistance of 100 tons or refusal. Refusal was 

considered to be less than 0.8 inch penetration in 10 hammer 

blows. 

4-3 
After the piling was driven, the existing embankment was excavated 

below the elevation of the highest tiebacks -- Figure 8. The 

piling was then cleaned and stud bolts were welded to the pile 

flanges. The stud bolts were later used to attach 4-inch by 8- 

inch treated wood lagging to the pile face. Exposed surfaces of 

the piling and bolts were protected by an application of "Tapecoat 

TC Mastic" -- Figure 9. 

The lagging was treated southern yellow pine timbers attached to 

the piling by threaded studs, steel plates, and nuts. There were 

approximately 2-inch gaps between the lagging through which the 

threaded studs protruded. The gap was spanned by metal plates and 

fastened in place with nuts. 

Treatment of the lagging consisted of a combination vacuum and 

pressure. The timber was subjected to a vacuum approaching 27 

inches of mercury for 30 minutes. The treatment solution was then 

pressure injected (140 psi) into the wood for 50 to 60 minutes. 

The solution consisted of 44.01 percent chromic oxide, 19.27 

percent cupric oxide and 36.72 percent arsenic pentoxide. A 3.27 

percent concentrate solution was used. 

Before the lagging was installed, a drainage pathway was placed 

between the wall and embankment. The pathway consisted of a layer 

of AMOCO 4553 fabric placed against the soil embankment and a 

layer of TENSAR "PWI" grid against the lagging -- Figures 6 and 

10. At the bottom of the wall, the pathway ended in a trough made 

of a cut section of corrugated plastic pipe. A collector system 

of 8-inch pipe was placed in the trough with outlet lines spaced 

at approximately 24 feet. The cavity behind the lagging was 

backfilled with the material previously excavated. The backfill 

was completed prior to testing, but in many cases failed to 

support the piling sufficiently during loading. Where the piling 

deflected too much the soil was removed and replaced with weak 

concrete. After placement of the drainage pathway, the lagging 

was installed beginning at the top of the piling. 

B. Tiebacks 

When the wall had been constructed down to the elevation of a 

tieback, a hole was drilled into rock and a steel tendon grouted 

in place. Each tieback was tested when the grout had reached 

sufficient strength. If the tieback tested acceptably, it was 

eventually locked off at 75 percent of design load. This lock-off 

load was chosen to permit some relaxation and movement of the 

retained embankment. The grout mixture contained Type III 

portland cement. 

Two tiebacks were placed in a bay midway between piles. The 

tiebacks were stacked vertically with alternating bays tied back. 

Toward the ends of the wall where the depth to rock decreased, one 

tieback on alternating bays was used. For a distance of 

approximately 110 feet (Station 234+80 to Station 23@+90), four 

4-4 
tiebacks on alternating bays were used. Where tiebacks failed, 

additional tiebacks were placed until tests indicated design 

restraint was achieved. 

In Figure 11, completed sections of the wall are shown. Double 

caps where supplementary tiebacks had to be installed may be noted 

in the foreground. At the top of the wall, a fence was erected to 

protect unwary pedestrians. 

A soldier beam and tieback schedule is shown in Table 1, page 43. 

Tiebacks are high-strength, in this case, multi-stranded, steel 

tendons anchored in rock at one end, stressed, and then anchored 

to the wall at the other end. 

is accomplished by drilling a 

The fixed anchorage or bond length 

hole (minimum of 10 feet) into 

competent rock and grouting the tendon into place. The hole must 

be clear of deleterious material and centralizers and spacers 

located so there is a minimum 

tendon. 

of l/2 inch grout cover on the 

The bond length is calculated by the equation: 

Lb = P/ (3.1416) (d) (tw) 

in which 

Lb = bond length (not less than 10 feet in solid rock) (feet), 

f; 

= design load for tieback (pounds), 

= diameter of the drill hole (inches), and 

t, = bond stress at the interface between rock and grout (psi). 

The unbonded length (Lf) is the portion of the tieback free to 

elongate elastically during stressing. This length is a minimum 

of 15 feet and is sufficiently long to insure that the bond length 

is formed in sound competent rock. Tiebacks were installed at 

angles varying from 10 degrees to 30 degrees from horizontal. A 

typical tieback cross section is shown in Figure 12. 

c. Instrumentation 

Several types of instrumentation were installed to monitor 

performance of the wall. Types of instrumentation consisted of 

equipment to stress and test a tieback, slope inclinometers, 

groundwater observation wells, tiltmeters, permanent load cells, 

and earth pressure meters. 

conducted. 

Optical surveys of the wall also were 

Equipment for stressing and testing the tiebacks consisted of a 

z;;;lic 

equipmint 

*jack for supplying.load, pressure gagesa::' monitoring 

dial gages for monitoring movement, 

such as jack stands, gage supports, etc. 

accessory 

Stressing and testing of the tiebacks will be discussed in the "TIEBACK TESTING" 

section. 

4-5 
Four permanent load cells were installed near Stations 235+90 

(Tieback 21) and 237+@4 (Tiebacks 43-44). At each location, two 

load cells were installed. One cell was on the upper or first 

tier tieback and one was on the lower or second tier tieback. 

These cells were used to monitor the short and long-term stresses 

on the tendons supporting the wall. 

Earth pressure meters were installed between the wall and earthen 

embankment. These meters are used to monitor pressure on the wall 

as opposed to the permanent load cells that monitor stress on the 

tendon supporting the wall. Nine meters were installed. Five 

were located in the center of the bay or midway between Piles 46 

and 47, Station 237+88. These five meters were numbered 1540, 

1447, 1541, 1449, and 1659, respectively, from the top of the 

wall. The other four meters were located as close as possible to 

Pile 47. They were numbered from the top 1444, 1615, 1542, and 

1614, respectively. The top meter on each row was placed 

approximately two feet below the top of the wall and the remaining 

meters were spaced at 2.0 to 2.5 foot intervals down the wall. In 

Figure 13, the top eight meters may be seen with the top two 

located between the third and fourth lagging from the top of the 

wall. 

As lagging installation proceeded down the wall, pressure meters 

were installed at the desired locations. The meters were loosely 

attached to the back side of the lagging with the monitoring 

cables exiting the open face of the wall. The cables were 

enclosed in plastic conduit and brought to a common monitoring 

point. After the meters were in place, the cavity behind the wall 

was backfilled with the previously excavated material and 

compacted with gasoline operated hand compactors. 

A total of seven slope inclinometers were installed to monitor 

horizontal earth movement. Five inclinometers were installed 

behind the wall near Stations 236+00, 237+00, 238+00, 239+00, and 

240+00. Location of the inclinometers ranged from 4 to 10 feet 

behind the wall. Two inclinometers were installed approximately 

85 feet right of centerline near Stations 236+00 and 237+00. 

Three ground-water observation wells were installed approximately 

35 feet right of centerline near Stations 237+50, 238+50, and 

239+50 l One observation well was installed approximately 30 feet 

left of centerline at Station 237+00. These wells, in conjunction 

with the slope inclinometer holes and observation wells installed 

during the earlier geotechnical investigation, permitted 

monitoring of the water table at the site. 

Tiltmeter plates were installed on the wall near Stations 236+00, 

237+00 and 238+50. These plates were installed on the wood 

lagging. Due to irregularities of the lagging and fluctuations of 

the characteristics of the wood related to changing moisture 

conditions, data obtained from these meters were not consistent. 

Location of instrumentation used to monitor the structure during 

and after construction of the wall is shown in Figure 14. 

4-6 
v. TIEBACK TESTING 

Each tieback was load tested by one of three types of tests. The 

three test types were creep tests , performance tests and proof 

tests. Creep tests and performance tests both essentially 

incrementally loaded and unloaded the tendon to 133 percent of the 

design load and monitored tendon elongation or movement. These 

two tests were conducted on a limited number of tiebacks and 

required a substantial amount of time. 

All tiebacks not tested by either creep or performance tests were 

proof tested. This test is of relatively short duration and 

consists of loading the tendon to 120 percent of the design load 

and maintaining that load five minutes. If creep movement during 

the five minutes is less than 0.03 inch and movement patterns are 

similar to adjacent tests the tieback is acceptable. If these 

criteria are not met, but the creep rate over a longer period of 

time was determined to be less than 0.08 inch per logarithmic 

cycle of time, the tieback was accepted. 

Four tiebacks were creep tested. This test required loading and 

unloading the tendon through gradually increasing load and time 

increments until 133 percent of the design load was reached. This 

load was maintained and movement observed for 300 minutes. An 

acceptable tieback performance was a creep rate less than 0.08 

inch per logarithmic cycle of time. 

Five selected tiebacks and 5 percent of all remaining tiebacks 

were performance tested. This test involved loading and unloading 

the tendon through gradually increasing load and time increments 

until 133 percent of the design load was reached. This load was 

maintained for 10 minutes after which time the test was 

discontinued if movement was less than 0.04 inch. If movement 

exceeded 0.04 inch the load was maintained for 60 minutes and the 

movement recorded Acceptable performance for these tests was 

(1) measured elastic movement exceeding 80 percent of the 

theoretical elongation of the unbonded tendon 

and 

(2) creep movement between 1 and 10 minutes less 

than 0.04 inches. 

Tiebacks failing criterion number 2 were accepted if the creep 

rate over 60 minutes of maximum loading was less than 0.08 inches 

per logarithm cycle of time. 

Equipment for testing the tiebacks may be seen in Figure 15. A 

hydraulic jack is affixed to the exposed tendons and load is 

applied. Resultant stress is monitored by gages not shown in the 

figure and the deflection is monitored with the dial gage mounted 

on the tripod. 

4-7 
VI. TEST DATA AND RESULTS 

A. Laboratory Data 

Laboratory tests were conducted on samples obtained at the slope 

inclinometer borings. Moisture content of samples tested ranged 

from 12.6 to 23.5 percent and averaged 17.7 percent. Specific 

gravity ranged from 2.67 to 2.81 with an average of 2.70. The 

material classified as A-6 or A-7-6 by the AASHTO system or CL by 

the Unified system. Results of index tests are contained in 

Table 2. 

A series of consolidated-undrained triaxial tests was performed. 

Results of these tests indicate an internal friction angle of 37 

degrees and a cohesion of 0. 

B. Lateral Movement8 

Slope inclinometers are identified as numbers 1, 4, 6, 8, 10, 11, 

and 12. The locations of Inclinometers 1 through 10 are behind 

the wall, approximately 40 feet right of the centerline, in order 

of ascending Stations 236+00, 237+00, 238+00, 239+00, and 240+00, 

respectively. Inclinometers 11 and 12 are located approximately 

85 feet right of the centerline at Stations 236++00 and 237+00, 

respectively. Inclinometers 1, 6, 8, and 10 were installed in the 

first week of March 1984, which was before excavation for wall 

construction began. Inclinometer 4 was installed May 5, 1984, 

after the wall was essentially complete. Inclinometers 11 and 12 

were installed June 18 and 11, respectively. 

Data obtained at Inclinometers 1 through 10 indicate little 

movement at depths greater than 4 or 5 feet. The exception to 

this is Inclinometer 1. This is approximately the location of 

monitoring instrumentation in place prior to corrective action. 

The sliding plane then was located approximately 25 feet below the 

surface. Data from Inclinometer 1 indicate a displacement of 0.5 

inch 47 days after installation of the inclinometer. This 

movement took place primarily along the existing sliding plane. 

Records indicate that final tiebacks near that location were 

stressed and locked off on May 11, 1984. After that date, little 

movement has been observed at that location -- Figure 16. 

All inclinometers near the wall indicate the greatest movement 

within 4 to 5 feet of the surface -- Figures 16 through 20. This 

is probably due to sloughing of the embankment after the material 

below the piling was excavated. The magnitude of this movement 

ranged from 0.5 inch to 1.5 inches. In Figures 21 through 23, 

movements at selected depths are plotted versus time and the dates 

of tieback lock off are noted. The top of the embankment was 

pulled back toward its original position when the tiebacks were 

stressed. In Figure 21, the movement toward the original position 

. illustrated for each inclinometer at approximately 125 to 150 

zys . This coincides with the locking off of tiebacks. At depths 

greater than 5 feet, embankment movement (with the exception of 

Inclinometer 1) was less than 0.3 inch. 

4-8 
Inclinometers 11 and 12 indicate maximum movements of 0.2 and 

0.8 inch, respectively. Movement at Inclinometer 11 is of such 

small magnitude that it is probably insignificant (Figure 24). It 

probably is erratic because it is approaching the limit of 

resolution of the instrument. 

occurring throughout 

At Inclinometer 12, movement is 

the entire depth of soil. As of March 1986 

0.8 inch of movement had occurred, but approximately 0.5 inch of 

movement occurred from February to November of 1985. Figure 25 

shows movement is continuing at this location, but the rate has 

decreased. 

As noted earlier, instrumentation intended to monitor tilt of the 

wall did not function properly. 

conducted to 

However, optical surveys were 

establish the initial position of the wall. 

Elevation of the top of the wall and plumb of the vertical face 

were established. No measurable changes have been observed. 

VII. PRESSURE DATA 

Earth pressure on the wall and retaining stress of the tiebacks 

were monitored with earth pressure meters behind the wall and 

permanent load cells on the tiebacks. 

used were high range 

The earth pressure meters 

meters 

meters was such that, 

(200 psi). Sensitivity of these 

except during testing and tieback lock off, 

pressure on the wall was too low to be monitored accurately. 

Initial readings were obtained prior to the backfilling operation. 

These readings were used as zero readings and subsequent 

were compared to them. 

readings 

During testing and lock off, pressure on the wall ranged from 0.0 

to approximately 

obtained 

10.0 psi at some locations. Pressure data 

27. 

from earth pressure meters are shown in Figures 26 and 

As seen in these figures, 

to be insignificant. 

long-term pressure on the wall 

appears The apparent negative pressure 

readings are a result of the low sensitivity of the 200 psi meters 

at the low existing pressure conditions. 

Permanent load cells were installed on Tiebacks 21 and 43-44. At 

both locations, lower 

tiebacks 

the higher (first tier) and 

instrumented. 

(second 

were 

tier) 

Tieback 21 is approximately located 

at Station 235+90 in a bay where four tiebacks were used. 

43-44 is Tieback 

approximately located at Station 237+64. 

location, 

At this 

only two tiebacks per bay were used. 

Lock-off loads on the tendons ranged from 68.8 kips to 132.0 kips. 

Loads on all tendons 

from 53.1 to 83.0 kips. 

decreased with time and eventually ranged 

Figure 28. 

Load cell data are plotted versus time in 

A. Water Table 

Water-table elevations were monitored with observation wells located at Stations 236+00 (45 feet left), 237+00 (30 feet left), 

237+50, and 238+50 and 239+50 (27 to 35 feet right) (see 

Figure 14). The water table fluctuated during construction of the 

4-9 
wall, but gradually rose after the wall was complete. Water-table 

depths are shown in Figure 29, with completion of the wall 

occurring at approximately 200 days. 

B. Durability 

Durability of the structure, 

and metal surfaces, 

primarily the exposed wood lagging 

was a major concern. Visual inspections and 

soundness checks indicated these wall components have not 

noticeably deteriorated during the first two years. 

VIII. PROBLEMS 

The most common problem associated with this project involved the 

inability of tiebacks to withstand test loading. Some of the 

reasons for tieback failure were the following: 

1. "Slick holes" 

in 

resulting from drilling an anchorage shaft 

the presence of water. Native rock at this site 

(Kope Formation shale) weathers rapidly in the presence 

of water. Bond between the grout and rock would be 

reduced in this case, thus permitting slipping of the 

tieback. 

2. Broken strands of the tendons accounted for several 

tieback failures. 

project, 

Of 121 tiebacks designed for the 

25 failed and were replaced or supplemented 

with additional tiebacks. 

Several tieback tests were discontinued due to excessive movement 

of the wall. In some cases, 

inches at only 50 percent of 

the piling deflected more than 4 

percent of the required 

design load or approximately 75 

resisting force. The solution to this 

problem was to excavate behind the piles and backfill with a weak 

concrete mix. 

Temporary delays were caused by failure of the welds on the 

threaded studs affixing the lagging to the piling and recalibration of the jack used to load the tiebacks. 

IX. CONCLUSIONS 

The tied-back wall, to the present, has performed well. Lock-off 

loads on the tiebacks were 75 percent of design load. Present 

loads are considerably less than lock-off. Pressure on the 

lagging appears to be insignificant. 

loading (p = 0.0528 kip/foot x h), 

Using the previously noted 

the wall at the pressure meter 

location was designed to support 7.33 pounds per square inch. 

Initial pressure on the lagging was about 50 percent of the design 

restraint with two meters briefly exceeding that valve (Figures 25 

and 26). This has 

locations. 

since dropped to nearly zero at all metered 

Pressure meters having better resolution would have 

been desirable for this application. However, the resolution of 

the meters is such that any significant pressures would have been 

measured. 

4-10 
Earth movement behind the wall has been controlled. The sliding 

failure existing prior to wall construction appears to have 

stabilized. In the vicinity of Station 236+00, movement along the 

sliding plane was observed until the tiebacks were stressed. 

Since that time, movement has been minimal. Below the wall at 

Slope Inclinometer 12 (85 feet right of Station 237+00), 

approximately 0.8 inch of lateral movement, Figure 25, was 

observed from October 10, 1984, to November 18, 1985. Movement is 

still occurring, but its rate has decreased. Surface slumping, 

due to excavation, was stabilized when the tiebacks were stressed. 

Since its completion, the wall has not moved or tilted. Optical 

surveys of the walls initial and more recent positions verify its 

stability. 

Perhaps the reduction of tieback stress and earth pressures may be 

explained by a combination of soil cohesion and relaxation of the 

wall components. Before the tiebacks were stressed, the 

embankment continued to slide and to slump where the soil had been 

excavated. When the tiebacks were stressed, those movements 

ceased. 

After stressing of the tiebacks the tieback tendons, piling, and 

lagging began to relax. This reduced the pressure on the wall 

components but restrained the embankment sufficiently to prevent 

it from moving along the sliding plane again. The cohesional 

component of shear strength (cementation) of the embankment may 

have prevented it from relaxing to the point of maintaining or 

increasing the original pressure on the wall. 

While the retained embankment is apparently stabilized,, the 

material below or in front of the wall may not be stabilized. The 

driving force on the embankment below the wall has been reduced, 

but movement continues. The rate of movement has decreased from 

0.05 inch per month to 0.009 inch per month, but monitoring will 

continue. If embankment movement should continue, alignment of 

the railroad and possibly the stability of the retained embankment 

could be adversely affected. The movement below the wall, 

25, 

Figure 

apparently justifies the assumption, Figure 5, of no soil 

resistance in.front of the wall. 

In general, the methods and materials used in construction of the 

wall appear acceptable. Problems, such as unacceptable pile 

deflection during testing, jack re-calibration, and poor stud bolt 

welding, caused relatively short delays. 

resulted from "slick holes" 

More significant delays 

and broken tendon strands. This led 

to placement of additional tiebacks. Overall, the construction 

progressed satisfactorily. 

It is anticipated that future design of similar structures will be 

less conservative. Due tieback wall has been constructed in 

Kentucky since the completion of the study structure. The second 

structures was designed with a safety factor of 1.0 and tested to 

133 percent of design load. 

4-11 





























REPORT NO. 6 

PROGRESS REPORT 

RETENTION SYSTEM MONITORING 

DEMONSTRATION PROJECT NO. 68 

NORTH STREET GRADE SEPARATION 

LIMA, OHIO 

For 

OHIO DEPARTMENT OF TRANSPORTATION 

INVESTIGATION BY 

THE H.C. NUTTING COMPANY 

CINCINNATI, OHIO 

1988 

TABLE OF CONTENTS 

I. DEMONSTRATION PROJECT NO. 68 . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6-l 

A. PROJECT LOCATION ..................................... 6-l 

B. SITE GEOLOGY ......................................... 6-l 

C. TIEDBACK RETAINING WALL PERFORMANCE STUDY ............ 6-l 

D. DESCRIPTION OF PERMANENT TIEDBACK RETAINING WALL 

COMPONENTS ........................................... 6-l 

E. TYPE AND LOCATION OF INSTRUMENTATION ................. 6-2 

II. PURPOSE OF INSTRUMENTATION . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6-3 

A. PROJECT INSTRUMENTATION .............................. 6-4 

B. INSTRUMENTATION .INSTALLATION ......................... 6-5 

C. INSTRUMENTATION DAMAGE DURING TIEBACK WALL 

CONSTRUCTION ......................................... 6-6 

D. INSTRUMENTATION MONITORING ........................... 6-7 

III. DATA ANALYSIS ,.......................................... 6-7 

A. SOIL PROFILE ......................................... 6-7 

B. DIGITILT INCLINOMETERS ............................... 6-8 

C. TIEBACK LOAD DISTRIBUTION ........................... 6-10 

D. SOLDIER BEAM NO. 93 AXIAL LOADS ..................... 6-13 

E. SOLDIER BEAM NO. 93 BENDING MOMENTS ................. 6-15 

F. TEMPERATURE READINGS AT SELECTED GAGE LOCATIONS ..... 6-16 

6-i 

A. PROJECT LOCATION 

The project (389-86) is located within the City of Lima, Allen 

County, Ohio. The project, ALL-S.R. 81-16.83 North Street Grade 

Separation, provides construction of an underpass below the 

railway serving the C & O/B & 0 Railroad Company and the N t W 

Railway Company. Reference is made to the location map attached 

with this report. 

B. SITE GEOLOGY 

The soils within the project site were deposited as glacial ground 

moraine during both the Illinoian and Wisconsin glaciation. The 

thickness of the soil cover above the bedrock varied from 

approximately 42 to 52 feet. The underlying bedrock is dolomite 

of the Tymochtee Member of the Monroe Group. 

The project soils are comprised of unsorted, unstratified mix of 

clay, silt and sand with a minor but variable percentage of gravel 

size material. The profile includes erratic seams and layers of 

sand and sand and gravel, discontinuous both horizontally and 

vertically. Generally, the sand and gravel is waterbearing as a 

result of surface water infiltration and accumulation in the more 

permeable granular zones. Typically, the soil possesses a very 

stiff to hard consistency, whereas the granular more permeable 

zones demonstrated a compactness varying between medium dense and 

very dense. 

c. TIEDBACK RETAINING WALL PERFORMANCE STUDY 

The North Street Grade Separation project in Lima, Ohio features 

the use of permanent soil anchors, soldier beams, and reinforced 

concrete facing wall in lieu of conventional cantilever retaining 

walls. Permanent tieback retention systems eliminate the need for 

temporary retention in congested areas thereby providing an 

economical alternative to the construction of standard cantilever 

walls. 

This study involves the monitoring of instrumentation to develop 

useful data for evaluating the long-term performance of the 

permanent tiedback retaining walls. The data provides documented 

magnitudes of stress within the components of the wall system 

which can be related to construction stages, field related 

conditions and design and construction effectiveness. 

D. DESCRIPTION OF PERMANENT TIEDBACK RETAINING WALL COMPONENTS 

Reference is made to shop drawings prepared by The Schnabel 

Foundation Company as well as the project plans prepared by the 

State of Ohio, Department of Transportation. Relevant drawings 

and plan sheets have been included in the appendix. 

6-1 
The instrumentation and monitoring was performed at wall unit No. 

71, soldier beam No. 93 and wall unit No. 22, soldier beam No. 24. 

Thirty-inch diameter drilled shafts were installed on 6 ft. 

centers. Between the bottom of the drilled shaft and the bottom 

of the wall footing the drilled shaft concrete was ODOT Class C 

mix which has a design compressive strength of 4000 psi. From the 

bottom of wall footing up to the top of the shaft, a lean concrete 

mix was utilized. The 7-day minimum compressive strength for the 

lean concrete was 1000 psi. Drilled shaft soldier beams included 

double channel C-15x33.9 with a clear interior spacing between 

channels of 14". The double channels were to be placed at plan 

location within the 30" diameter drilled shaft. The steel 

channels were ASTM Designation A-36 Grade Steel. 

As the excavation progressed, wood lagging was installed behind 

the steel beam flanges after removing drilled shaft lean concrete 

at the exterior face flanges. The lagging consisted of nominal 3" 

thick rough cut mixed hard woods which were untreated. 

The permanent tiebacks were installed at various stages of the 

excavation in accordance with plan requirements. The drill holes 

for the anchors were 12" diameter and drilled using a hollow stem 

auger rig. The tieback tendon was placed within the hollow stem 

and carried to the end of the drill hole by means of a closure 

device at the lower end of the drill stem. The closure device was 

disengaged and the auger stem withdrawn during the grouting 

process. The grout composition consisted of a 9 bag mix (Portland 

cement Type I), sand and fly ash. 

The tiebacks included a double corrosion protection system. The 

tiebacks were l-1/4" diameter Dywidag thread bars 

within a 

encapsulated 

corrugated sheathing 

The l-1/4" 

as supplied by the manufacturer. 

diameter Dywidag bars are 150 ksi grade steel 

conforming to ASTM A-722. 

After the basic retention system was completed, a reinforced 

concrete facing wall was constructed. The facing wall was 

anchored to the flanges of the double channel soldier beams by 

studs welded to the channels. A2 ft. wide 

constructed at the base of the facing wall. 

strip footing was 

E. TYPE AND LOCATION OF INSTRUMENTATION 

The instrumentation program included the use of 

inclinometer casing, 

digitilt 

load cells and vibrating wire strain gages. 

The instrumentation was installed at two wall sections; one at 

soldier beam No. 93, wall unit No. 

wall unit No. 22. 

71 and at soldier beam No. 24, 

The instrumentation at soldier beam No. 

inclinometer casing full length within 

24 included digitilt 

the soldier beam. The 

casing was attached on the interior side of the south channel 

flange. Instrumentation at soldier beam No. 24 also included load 

6-2 
cells on both the first and second level anchors at the anchor 

head. Load cells were installed after the individual anchors were 

load tested. 

Instrumentation at soldier beam No. 24 also included strain gages 

installed directly on the two l-1/4" diameter Dywidag bar 

tiebacks. Gages were placed at the midpoint of the unbonded 

anchor length and at four locations within the bonded anchor 

length. The gages installed within the unbonded anchor length 

were 15 ft. beyond the anchor head. The gages within the bonded 

anchor length were installed at distances of 2 ft., 6 ft., 10 ft. 

and 18 ft. beyond the beginning of the bonded anchor length. 

Figure 4 illustrates the typical strain gage and load cell 

locations for the instrumented tiebacks. As noted, a 30 ft. 

unbonded anchor length and a 20 ft. bonded anchor length were 

typical for instrumented tieback anchor installations. 

The instrumentation at soldier beam No. 93 was similar to that for 

soldier beam No. 24, however, additional instrumentation included 

strain gages placed on the double channel soldier beams 

(2xClSx33.9). Four horizontal sections through the soldier beams 

were chosen for strain gage instrumentation. Gages were placed at 

the same elevation as the level one tieback. This section 

included four strain gages, one each on the exterior flanges of 

the soldier beam channels. The second section was chosen at a 

location midway between the second level tieback and the bottom of 

facing wall footing. Section 2 included four strain gages, one 

each on the exterior face of the channel flanges. Section 3 was 

chosen at a location corresponding to the bottom of the facing 

wall footing elevation. In addition to the four strain gages on 

the channel flanges, two strain gages were placed on the center 

line of each channel web. These gages were placed on the face of 

the web between the protruding flanges. The instrumentation at 

Section 4 was located 5 ft. above the bottom of the soldier beam. 

The instrumentation at this section included two strain gages on 

the centerline of each channel web similar to the installation at 

Section 3. This section did not include strain gage 

instrumentation on the channel flanges. 

Figures 1 thru 4 illustrate strain gage, load cell and digitilt 

inclinometer location details as well as the construction details 

of the tiedback wall. 

II. PURPOSE OF INSTRUMENTATION 

The primary purpose of the instrumentation program is monitoring 

the long-term performance of the permanent soil tiebacks. A 

secondary benefit of the instrumentation program is the monitoring 

and evaluation of the overall tiedback wall performance on a 

short-term and long-term basis. 

6-3 
The primary components of the tieback retention system constructed 

for the North Street Grade Separation project are the individual 

soldier beams and tiebacks. The evaluation of tieback system 

performance relative to these components addressed four key 

elements: 

1. The rate of soldier beam deflection and tieback load 

changes, 

2. The magnitude of soldier beam deflection and tieback 

loads, 

3. The direction of soldier beam deflection, and 

4. The specific location and magnitude of deflection along 

the length of the soldier beam and tieback load 

development along the anchor length. 

Digitilt inclinometers, load cells, and strain gages were utilized 

to obtain data for evaluation of the four items listed above. 

Strain gage installation on the soldier beams permits analysis of 

both axial and bending stress at various locations along the 

soldier beam length. This instrumentation, at least to some 

extent, can be used to evaluate the design performance of the 

soldier beam components To a lesser extent, this data can be 

correlated with the other tiedback wall component data to permit a 

more thorough evaluation of the overall tiedback retaining wall 

system performance. 

The relationship between the various instrumentation provides some 

refinement in overall performance evaluation, especially with 

respect to the load cell data and the data from the, strain gages mounted on the tiebacks. Further, the deflection data obtained 

from the digitilt inclinometer monitoring permits interpretation 

of tieback load changes. Some additional evaluation benefit is 

derived through the relationship between the strain gage data on 

the soldier beam channels and the other instrumentation data. 

A. PROJECT INSTRUMRRTATION 

The following is a list of instrumentation and monitoring 

equipment utilized for this study. 

SINCO: Slope Indicator Company 

Digitilt Inclinometer Sensor Probe Model 50325 

Digitilt Inclinometer Indicator Model 50306 

Digitilt Inclinometer Casing - ABS Plastic 

1.90" O.D.xl.50" I.D. 

6-4 
IRAD GAGE: 

Vibrating Wire Hollow Load Cell, Model VH-150 

Spot-Welded Vibrating Wire Strain Gages, Model SM-2W with 

thermistors. 

Vibrating Wire Manual Switching Stations, Model TB-12s and 

Model TB-24s 

IRAD/KLEIN Smart reader M Model SR-11 

B. INSTRUMENTATION INSTALLATION 

The digitilt inclinometer casing was installed by the specialty 

contractor, Schnabel Foundation Company, by temporarily fastening 

the casing with metal straps to the web and flange of the soldier 

beam channel. After installation of the soldier beam channels 

into the drilled shaft, the drilled shaft concrete secured the 

casing permanently into place within the soldier beam. The 

digitilt inclinometer casing extended full depth within each of 

the monitored soldier beam locations. Figures 1 thru 3 illustrate 

the location of the digitilt inclinometer casing and the reference 

elevations for the subsequent monitoring readings. 

Strain gage installation on the l-1/4" Dywidag thread bars was 

performed by a representative of IRAD GAGE. The installation was 

performed at the Dywidag Systems International USA, Inc. 

production plant in LeMont, Illinois. Installation of the strain 

gages at the plant was necessary due to the fact that the tiebacks 

were required to have a double corrosion protection system. The 

gages I therefore, had to be installed prior to encapsulating the 

Dywidag bars within the corrugated sheathing. 

The strain gages were installed on the tiebacks at the locations 

illustrated in Figure 4. The strain gages were spot-welded to the 

Dywidag bar after standard surface preparation was completed. The 

strain gages were protected by a standard coil/magnet cover. Each 

of the strain gages included a thermistor for monitoring 

temperature at selected gage locations. Since the tiebacks are 

tension elements, the initial strain gage wire tension was 

adjusted in order to increase the tensile strain range of the 

gage ' 

In order to facilitate installation of the instrumented tiebacks 

during construction, it was necessary to cut each of the strain 

gage read-out cables so that the cables did not extend beyond the 

head end of the tieback within the hollow stem augers. Each of 

the cables was labeled with a corresponding reference to a gage 

location. After installation of the tiebacks, it was necessary to 

splice each of the read-out cables so that there was sufficient 

length of cable to reach the final location of the switching 

stations. 

6-5 
Strain gage installation on the double channel beams for soldier 

beam No. 93 were performed by a representative of IRAD GAGE. This 

installation was made at the job site within a secured area. 

Again, the gages and coil/magnet covers were spot-welded to the 

channel beams after standard surface preparation. Figures 2 and 3 

illustrate the locations of the strain gages installed on the 

channel beams. 

Efforts were made to protect the instrumentation on the channel 

beams from potential damage during installation and concrete 

placement. Small steel plate sections were welded over each 

strain gage cover. The read-out cables from the strain gages were 

placed within 3/4" diameter PVC pipe which was extended to the top 

of the soldier beams. The PVC pipe and limited sections of 

exposed read-out cable were secured with liquid adhesive. 

The load cell was installed on each of the tiebacks after the load 

testing at each location had been completed. The load cell was 

sandwiched between two anchor bearing plates, the tieback 

tensioned to lock-off load and the anchor nut then secured in 

place. The load cells were installed by the specialty contractor. 

The lead cables from the load cells were protected in 3/4" 

diameter PVC pipe which was attached to the wood lagging to 

inhibit damage during construction. 

During the later stages of wall construction, the digitilt 

inclinometer casing which protruded above the wall beam cap was 

protected within a permanent steel casing with a lockable hinged 

cap. The vibrating wire switching stations which were utilized to 

facilitate strain gage readings were placed in standard concrete 

electrical pull boxes immediately behind the wall location of 

soldier beam Nos. 24 and 93. The pull boxes and the steel casing 

over the digitilt inclinometer casing provides reasonable 

protection against vandalism and accidental damage. 

c. INSTRUMENTATION DAMAGE DURING TIEBACK WALL CONSTRUCTION 

During the early stages of tieback wall construction, two gages 

which were located on channel flanges were damaged. These two 

gage locations were at Section 1 on the south flanges of the 

channel beams facing the excavation. These gages were damaged 

during the removal of the drilled shaft lean concrete along the 

channel flanges in order to provide clearance for the lagging 

installation and reinforced concrete facing wall construction. 

Eight of the read-out cables for the strain gage and load cell 

instrumentation were damaged by the excavation equipment at 

soldier beam No. 93. However, the read-out cables were repaired 

by splicing additional length of cable after removing the damaged 

section of cable. 

As indicated in Tables I thru V, a number of the strain gages were 

non-functioning at some point in time after initial readings were 

6-6 
made. The non-functioning gages include gage Nos. 20, 35, 42,45 

and 48. Gage No. 43 could have been over-stressed during tieback 

anchor stressing to lock-off load, however, this is not certain. 

The reason that the other gages became non-functional was not able 

to be determined. 

D. INSTRUMBNTATION MONI!CORING 

The instrumentation monitoring commenced July 16, 1987. The 

soldier beam installation had been completed at soldier beam 

location Nos. 24 and 93. initial excavation along the wall had 

also been made and the wood lagging operations were in progress. 

The initial monitoring of the instrumentation established the base 

readings for referencing subsequent instrumentation data. 

For all gages other than the load cells, initial instrumentation 

readings were not possible until after field installation of the 

tieback and soldier beam components. As an example, the 

instrumented tiebacks and channel beams were subjected to 

different strains in the installed position of these elements as 

compared to their pre-installation strains. 

The instrumentation monitoring reflected various stages of 

tiedback retaining wall construction progress. A log of field 

conditions on the dates of instrumentation readings was maintained 

to provide relevant information for evaluating wall system 

performance. Reference is made to Figure 5. The log of field 

conditions documents the excavation elevation adjacent to the 

wall, stage of wall construction, and other data pertinent to wall 

system performance. 

III. DATA ANALYSIS 

The instrumentation data is illustrated graphically in Figures 5 

thru 11 and Graphs 1 thru 18. Discussion of observations and data 

analysis covering the initial 4-l/2 months of tieback retaining 

wall performance is provided below. 

A. SOIL PROFILE 

In addition to the general subsurface exploration performed in 

conjunction with the project, a single test boring was performed 

in the proximity of soldier beam No. 93 to develop specific soil 

data at an instrumented wall section. The test boring and 

associated laboratory tests were performed for, and funded by, the 

Schnabel Foundation Company. 

The laboratory test results were evaluated in terms of the 

suitability of the encountered soils for installation of permanent 

soil tiebacks. Specifically, analysis determined whether project 

cohesive soils are creep susceptible based on generally accepted 

criteria. The first criteria considered the Consistency Index 

(Ic) which is the difference between the soil liquid limit and the 

6-7 
in situ moisture content with respect to the difference between 

the liquid limit and the plastic limit of the soil. The test 

results indicated a consistency index varying between 1 and 1.3. 

This range exceeds 0.9, therefore, indicating non-creep 

susceptible cohesive soils. Other generally accepted criteria 

requires that in situ moisture content be at or below the plastic 

limit of the soil. Project soils satisfy this criteria as well. 

Unconfined compressive strengths vary between 2.35 and 4.22 tons 

per sq. ft. which exceeds the minimum 1 ton per sq. ft. criteria 

for non-creep susceptible cohesive soils. Effective strength 

parameters of the soil were determined by performing consolidated 

undrained triaxial compression tests with pore pressure 

measurements. The testing indicated an angle of internal friction 

of 35O and a cohesion value of 340 lbs. per sq. ft. 

B. DIGITILT INCLINOMETERS 

The digitilt inclinometer determined deflection perpendicular to 

the wall (north-south primary orientation) as well as parallel to 

the wall (east-west secondary orientation) at soldier beam Nos. 24 

and 93. On two occasions during the monitoring process, it was 

not possible to determine secondary deflection due to 

malfunctioning of the sensors within the probe. 

Figures 6 thru 11 provide graphic plots of the deflection data. 

The plots also include the locations of the two tieback levels, 

the bottom of wall footing and other relevant data. It should be 

noted th,at the top of casing elevation shown on the plots 

represents the elevation at the cable clip on the inclinometer 

assembly. This elevation also serves as the zero depth basis. 

According to the equipment manufacturer, the digitilt inclinometer 

data has an overall accuracy of 0.3" per 100 ft., of inclinometer 

casing length. Based on the inclinometer casing lengths for this 

project, overall accuracy is on the order of 0.1 to 0.15". 

The inclinometer data obtained in the primary orientation at 

soldier beam No. 24 indicates essentially no deflection occurred 

below the level of the first tieback anchor subsequent to the 

first two readings. Reading No. 1 was made after tensioning of 

the first level tieback and the data illustrates that the top of 

the soldier beam was pulled toward the retained soil approximately 

0.25". As as results, the point of soldier beam rotation relative 

to the base readings occurred at a depth of 13 ft. or 

approximately elevation 863.5. The second reading was made after 

the second level tieback had been tensioned. The excavation 

locally adjacent to the wall had also been taken to the bottom of 

wall footing elevation. This and subsequent readings indicated 

that below the first tieback level deflection did not exceed 

approximately 0.1". The portion of soldier beam above the first 

tieback level behaved similarly to a cantilever beam 

demonstrated by deflection data from the subsequent readings. T:: 

subsequent readings indicated that the top of soldier beam 

deflected toward the excavation on the order of 0.3" relative to 

6-8 
data developed during reading No. 1. As clearly demonstrated by 

the later readings, the upper portion of the soldier beam has 

rotated about a point corresponding to the first tieback level. 

Secondary orientation inclinometer data for soldier beam No. 24 

indicated initial movement of the soldier beam in an eastward 

direction as reaction to the tensioning of the first level 

tieback. The upper half of the soldier beam deflected 

approximately 0.1" to the east. Following lock-off of the second 

level tieback, 

initial location. 

the soldier beam returned essentially to its 

stabilization of 

Subsequent to August 5, 1987, there was a 

soldier beam deflection in an east-west 

direction. The upper half of the soldier beam experienced 

deflections within the range of zero to 0.1" west of its initial 

location. It is probable that completion of the concrete 

wall by August 21, 

facing 

1987, accounted for stabilization of movement 

in the east-west direction. 

Data developed at soldier beam No. 93 in the primary axis 

indicated deflection similar to that at soldier beam No. 24. Data 

from reading No. 1 showed that the top of soldier beam was pulled 

into the retained soil approximately 0.15" as a direct result of 

tensioning the first level tieback. Reading No. 2 was made 

immediately following tensioning of the second level tieback. 

This reading clearly indicated the impact of tensioning the second 

level tieback relative to soldier beam deflection. Immediate 

soldier beam deflection was on the order of 0.05" within the local 

area of the second level tieback with magnitude of deflection 

diminishing toward the top and bottom of the soldier beam. The 

first two readings indicated that the point of rotation was at an 

approximate depth of 12 ft. or elevation 865. Reading No. 3 was 

taken following local excavation to the wall footing level. In 

response, the soldier beam deflected toward the excavation between 

0.05" and 0.1". Subsequent readings indicated a slight increase 

in deflection toward the excavation with a maximum value near the 

level of the second tieback of 0.2". The soldier beam deflection 

demonstrated a broad parabolic deflection distribution. It would 

also appear that the soldier beam deflection has stabilized since 

the end of September 1987. Data at soldier beam No. 93 in the 

secondary axis illustrated a distinct reverse curvature of the 

soldier beam at a depth of approximately 18 ft. It is probable 

that this phenomenon is the result of excavation equipment contact 

during excavation. 

following 

Reading 2 provided deflection data immediately 

tensioning of the second level tieback. Tensioning of 

the second level tieback reduced the localized protrusion of the 

soldier beam eastward at the 18 ft. depth level. At this point, 

the maximum eastward deflection was reduced to 0.15" with deflection decreasing to initial soldier beam location at the top 

and bottom. Reading No. 3 demonstrated the effects of excavation 

to the bottom of wall footing elevation. Deflection increased to 

the east as much as 0.15" within the portion of soldier beam below 

the level of the first tieback. Subsequent readings indicated a 

slight soldier beam shift to the west. Further, it is apparent 

6-9 
that the portion of soldier beam above the first level tieback has 

behaved as a cantilever beam. Completion of the reinforced 

concrete facing wall for the retaining system stabilized east-west 

deflection, as would be expected. 

Comparison of the digitilt inclinometer data developed at soldier 

beam No. 24 and 93 within the primary orientation revealed some 

differences in the deflected shapes of the soldier beams. The 

later readings at soldier beam No. 24 indicated an S-shaped 

deflection distribution. Maximum deflection toward the excavation 

occurred between the second level tieback and the bottom of 

footing. Magnitude of deflection was on the order of 0.1". The 

upper section of the soldier beam behaved similarly to a 

cantilever beam bending about the first level tieback. The data 

developed at soldier beam No. 93 indicated a parabolic shaped 

deflection distribution. At this location, maximum deflection 

toward the excavation occurred within the vicinity of the second 

level tieback having a magnitude of 0.2". At this location, there 

was no reverse curvature in the deflected shape of the soldier 

beam at the first tieback level. 

The digitilt inclinometer readings to date have revealed 

relatively minor soldier beam deflection. The magnitude of 

soldier beam deflection is consistent with or slightly less than 

what would generally be expected for large diameter, straight 

shaft permanent soil anchors in cohesive soils. 

c. TIEBACK LOAD DISTRIBUTION 

Instrumentation of the tiebacks included a load cell installed at 

the anchor head for both tieback levels at soldier beam Nos. 24 

and 93. Each of these tiebacks also included vibrating wire 

strain gage instrumentation at five locations along the anchor 

length. The strain gage locations are shown on Figure 4. The 

strain gage data was converted to stress utilizing the theory of 

elasticity. The resulting stress was used to determine the 

tieback loads at the specific strain gage locations. The load 

cell data was used to determine tieback loads at the anchor head 

utilizing calibration charts and equations provided by the load 

cell manufacturer. 

The load cell and tieback forces are shown graphically with 

respect to time since initial tieback lock-off on Graphs 1 thru 4. 

The design and lock-off load for these tiebacks was 80 kips. The 

data indicates that the tiebacks have been locked-off at a force 

less than 80 kips. Initial tieback loads after lock-off varied 

between 53 and 70 kips. The low lock-off loads are more than 

likely attributable to the lock-off load operation using the 

stressing jack, limited accuracy of the pressure gage on the 

stressing jack, seating of the anchorage at the bearing plate and 

other construction related factors. Data obtained immediately 

following lock-off of the second level tieback anchor at soldier 

6-10 
beam No. 93 showed that the design load was not achieved at lockoff. The initial reading was 70 kips as compared to the 80 kip 

design load. 

Data from three of the four load cells indicated tieback load loss 

of 1% or less between the initial readings taken late July or 

early August, 1987 through the early part of December 1987. These 

percentages are based on the difference in load over the 

monitoring period to date as compared to initial anchor loads. 

The load cell at soldier beam No. 24, first level tieback, 

indicated an approximate 6% loss of tieback anchor load within the 

monitoring period to date. A substantial loss of load was 

recorded between the readings taken August 14 and 21, 1987, 

approximately 7 kips. A good portion of this load loss was 

recovered based on readings taken the following week. Beyond this 

point the load cell indicated stabilization of tieback load. Less 

than 2% load loss occurred within the final 3-l/2 months of the 

monitoring period through December 1987. It would appear that the 

7 kip load loss in August represented shift of anchor bond 

capacity further back on the anchor length. 

In addition to the load cells, the strain gages mounted on the 

tieback bars were utilized to evaluate load transfer with time 

along the length of the anchor. Figure 4 illustrates the 

locations of the strain gages along the anchor length. The strain 

gages located at the approximate mid-point of the unbonded length 

should theoretically yield loads which agree with the tieback 

loads indicated by the load cells at the anchor head. At the 

beginning of the monitoring period, the strain gage readings on 

the tiebacks revealed some difference in tieback load as compared 

to the load cell data. There was a gradual increase in anchor 

load at the No. 1 strain gages over the length of the monitoring 

period. The December 1987 readings disclosed an approximate 1 kip 

or less load difference between the data developed from the load 

cells and No. 1 strain gages. As noted at soldier beam No. 24, 

first level tieback, the No. 1 strain gage was not functioning. 

It is probable that there was some initial friction between the 

tieback anchor bars and the smooth sheaths in the unbonded anchor 

length. With time, the unbonded zone was fully mobilized. Tieback loads at the No. 1 strain gage locations increased 3.4 to 

10.9 kips over the first 4-l/2 months of monitoring. 

Strain gage Nos. 2 thru 5 on each of the instrumented tiebacks are 

located within the bond length of these anchors. It was intended 

that the strain gage data provide tieback anchor load distribution 

along the anchor length as well as anchor load transfer with time. 

The gages did show distribution of anchor load within the bonded 

anchor length of the tieback. A significant portion of the anchor 

load, however, was developed within the anchor zone preceding the 

bond length. Undoubtedly, the anchor grout immediately preceding 

the bond zone is in compression and provides significant load 

6-11 
capacity as a result of the anchor grout/soil bond. To illustrate 

this point, the No. 2 gages revealed initial anchor loads of 

approximately 21 to 36% of the tieback load indicated by the 

corresponding load cells. The December 1987 readings at the No. 2 

gage locations demonstrated a similar load range, about 31 to 33% 

of the tieback loads indicated by the load cells. As noted, the 

No. 2 gages are located only 2 ft. beyond the beginning of the 

bonded anchor length. Therefore, more than 67% of the anchor load 

capacity is developed within the length of anchor preceding the 

No. 2 gage locations. 

It is interesting to note that No. 2 strain gage on the first 

level tieback of soldier beam No. 93 decreased 4.9 kips upon 

tensioning of the second level tieback. This response illustrates 

the inter-dependence of the retaining system components in an 

effort to reach stress equilibrium. 

The data developed at strain gage locations 3, 4, and 5 were also 

evaluated relative to percentage of tieback load indicated by the 

load cells. At the No. 3 gage locations, initial readings 

suggested 3.8 to 5.2% of anchor load distribution. The December 

1987 gage readings increased to approximately 15% of anchor load 

distribution. At the No. 4 gage locations, initial readings 

indicated 1.2 to 6.7% of anchor load distribution as compared to 

the range of 7.7 to 13.6% of anchor load distribution for the 

December 1987 readings. The No. 5 strain gage locations indicated 

initial tieback load percentages of 0.5 to 2.4%. The December 

1987 readings indicated 5.4 to 5.7% of anchor load distribution. 

It should be noted that there were a number of non-functioning 

strain gages on the instrumented tiebacks. Further gage Nos. 2,3, 

and 5, at the second anchor level of soldier beam No. 24, yielded 

peculiar data. In fact, gages Nos. 3 and 5 suggested compression 

loads. Anchor tensioning may have introduced local bending 

strains within the bar tieback which misrepresents actual tieback 

loads. At low axial strains, gages mounted on one side of the 

tiebacks may reflect compression bending strain. If the data is 

reviewed in terms of increase in anchor load, the strain gage data 

at these locations have some value. 

The instrumentation data shows tieback load has remained 

relatively stable at the anchor head. However, the tiebacks have 

demonstrated some increase in anchor load distribution at the 

various locations along the unbonded and bonded anchor length. 

Generally, the first 3 to 5 weeks after tieback lock-off has 

demonstrated most of the load increase at the strain gage 

locations. The increase in stress along the bonded anchor length 

is probably a function of many factors including excavation 

adjacent to the wall, tieback tensioning at adjacent and lower 

level anchors, creep between the grout anchor zone and the 

surrounding soil and the time dependent stress relaxation of the 

prestressed steel tiebacks. Literature suggests a 3 to 4% tieback 

6-12 
load loss at the anchor head can be expected i the 4-l/2 month 

monitoring period due entirely to the stress relaxation of the 

prestressed steel. 

D. SOLDIESRBgAWNO. 93AXIALLOADS 

The strain gage instrumentation placed on the double channel beams 

(soldier beam No. 93) provided data to determine axial and bending 

stress in these members at specific locations along the soldier 

beam length. Axial loads were determined for both the steel 

channels as well as the composite section of steel beams and 

concrete of the surrounding drilled pier. Sections 3 and 4 

included the full drilled pier concrete area, whereas Section 2 

analysis accounted for the smaller concrete area as a result of 

concrete removal to install the reinforced concrete facing wall 

and wood lagging. Reference is made to Graphs 5 thru 11; all 

present the results of strain gage instrumentation on soldier beam 

No. 93. 

Readings taken early in the monitoring program indicated that gage 

3.2 was yielding erroneous data. It is probable that the 

excavation equipment or the concrete removal process at the face 

of the channel flanges locally deformed the channel flange which 

resulted in distortion of the data. The data has been presented 

in two ways. The first considers the actual data from gage 3.2. 

The other procedure provides a theoretical correction of gage 3.2. 

The theoretical correction of gage 3.2 was determined by 

developing a linear strain distribution north to south through the 

east channel beam. Gages 3.4 and 3.6 were utilized for this 

purpose. 

As noted previously, the two strain gages on the south flanges of 

the channel beams at Section 1 of soldier beam No. 93 were damaged 

during concrete removal for lagging installation. Therefore, the 

strain gage data developed at Section 1 consisted of only the 

north flanges of the channel beams. These gages indicated there 

were low compression strains in these flanges during the early 

monitoring period. After the middle of August 1987, there was a 

transition to tensile strain. Obviously, the data manifests 

bending of the upper portion of the soldier beam as a result of 

progressive mobilization of earth pressure acting on the wall. 

The developed stress within the channel beams is on the order of 

1% of the allowable steel stress. 

Strain gage instrumentation at Section 2 indicated an average 

stress within the steel channel beams between 3 and 4% of 

allowable within the later part of the monitoring period. 

Converting strains to axial load for both the steel channel beams 

and the composite section revealed low initial axial loads with 

dramatic increase of axial loads within the first month of 

monitoring. A 6.2 kip increase in axial load occurred within the 

composite section as a result of tensioning the second level 

tieback on July 28, 1987. Another large increase in axial load 

6-13 
occurred during the following week as the excavation was extended 

to bottom of wall footing elevation. The axial load on the 

composite section showed a 10.5 kip increase due to the reduction 

in drilled pier shaft friction related to the excavation. 

Monitoring readings taken August 14, 1987, revealed the greatest 

jump in soldier beam axial load, approximately 27.5 kips. It is 

probable that this jump occurred due to concrete removal at the 

south flanges of the steel channel beams relative to lagging 

installation. A hoe ram and pneumatic hammer were used to remove 

the concrete. This operation produced considerable vibration of 

the soldier beam and permitted axial load transfer to greater 

depths on the soldier beam. 

There was some continued increase in soldier beam load at Section 

2 until the middle of September and then a gradual decrease in 

axial load through the end of the monitoring period to date. 

Maximum axial loads in the composite section reached 55.4 kips. 

The vertical component of axial load due to the tieback tensioning 

at the first and second level amounts to 46 kips. The weight of 

the pier above Section 2 is approximately 13 kips. A total 

vertical load of 59 kips is theoretically developed. This does 

not include the weight of the reinforced concrete facing wall 

which at least partially is carried by the soldier beam. 

Therefore, most of the load produced by the various wall 

components has been carried to the level of the Section 2 gages. 

The data also suggests there is an instantaneous as well as time 

dependent response to axial loads produced by tieback anchor 

tensioning, dead load of the wall system components, excavation 

adjacent to the wall, and general construction procedures. 

Discussion related to the Section 3 strain gages will be limited 

to the data based on the rectification of gage 3.2. The 

calculated stress within the steel channel beams at this section 

was on the order of l/2 to 1% of allowable stress and on the same 

order of magnitude as Sections 1 and 2. The Section 3 data did 

not indicate the instantaneous response to various construction 

operations as demonstrated at Section 2. There was a time 

dependent transfer of soldier beam axial load to the Section 3 

location. However, axial load transfer was considerably more 

gradual than encountered at Section 2. Further, the composite 

section axial loads were on the order of 30% of the loads 

developed at Section 2 during the later stages of instrumentation 

monitoring. In fact, there was only an approximate 10 kip 

increase in axial load on the composite section between the 

initial readings and the December 1987 readings. 

The maximum axial load was approximately 17 kips in December as 

compared to a computed 63 kips of axial load. The 63 kips 

includes the vertical component of the tieback loads and weight of 

the soldier beam pier but excludes any partial load developed due 

to the facing wall. Therefore, the majority of the soldier beam 

axial load was supported through drilled pier shaft friction 

within the 5 ft. length of soldier beam between the Section 2 and 

6-14 
Section 3 gages. It should be noted that the finished pavement 

grade is approximately elevation 854 adjacent to the wall or 

approximately 1 ft. below the Section 2 gages. The reinforced 

concrete facing wall footing probably supports some of the axial 

load transmitted to the soldier beam considering the integral 

fabrication of the reinforced concrete facing wall to the soldier 

beam units. 

Section 4 strain gage instrumentation included placement of strain 

gages on the centerline of the webs on each channel beam. The 

developed strains were extremely low. The data indicates that 

gage 4.6 experienced a much greater strain rate change than gage 

4.5. This difference may be due to concentration of local bending 

strains at the location of Section 4. Averaging the strain data 

at Section 4 yields composite section axial loads up to a maximum 

of 3 kips. The computed axial load at Section 4 is 70 kips and is 

the summation of the vertical component of the tiebacks and the 

weight of the soldier beam concrete and steel. It is interesting 

to note that essentially all of the strain at Section 4 has been 

carried by the east channel beam, especially as indicated by the 

data recorded in November and December 1987. 

Most of the axial load on the soldier beam has been transmitted to 

Section 2, just above final grade elevation adjacent to the wall. 

The net axial load quickly diminished with increased depth of 

soldier beam embedment below finished grade. The stress developed 

in the steel channel beams is less than 4% of allowable based on 

the strain gage readings. Additionally, if the maximum vertical 

load as recorded at Section 2 (55.4 kips) was carried entirely by 

the channel beams, neglecting the surrounding concrete, stress 

levels would only be approximately 12% of allowable. 

E. SOLDIER BEAM NO. 93 BENDING MOMENTS 

The strain gage instrumentation at soldier beam No. 93, Sections 2 

and 3, was used to determine bending moments in the soldier beam. 

At Section 2 the bending moments were determined for the steel 

channel beams only. The complications in determining composite 

area bending moments at this section were prohibitive considering 

the fact that a portion of the concrete had been removed for 

lagging installation and accommodation of the facing wall. 

Analysis of Section 3 data permitted determination of bending 

moment in the steel channel beams as well as the composite section 

considering an untracked concrete section. Reference is made to 

Graphs 12 thru 16. 

Similar to the determination of axial soldier beam loads, gage 3.2 

was theoretically adjusted as described in the previous section of 

this report. Discussion of bending moments at Section 3 has been 

limited to the data which considers the adjustment made to gage 

3.2. 

6-15 
At both Sections 2 and 3, the bending moments were higher in the Y 

axis (direction parallel to the wall) than bending moments in theX axis (direction perpendicular to the wall). Another peculiar 

results indicated by the data is that the individual channel beam 

bending moments in the X axis were in opposite directions. In 

other words, the west channel beam indicated compression strains 

in the south flange while the east channel beam indicated tensile 

strains in the south channel flange. It is apparent that soldier 

beam torsion has considerably affected the strain gage readings. 

Another possible factor contributing to the peculiar strain gage 

data is that the bond between the steel channel beams and the 

surrounding concrete is questionable due to the ram hoe and 

pneumatic hammer operation during concrete removal along the south 

flanges of soldier beam No. 93. 

The data suggests that bending in the Y axis became stabilized 

after the reinforced concrete facing wall was completed following 

the August 21, 1987, readings. 

The developed strain gage data indicates that bending moments 

within the soldier beam were extremely low relative to the 

structural capability of the steel channel beams and composite 

soldier beam section. 

F. TRMPERATDRE READINGS AT SELECTED GAGE LOCATIONS 

Thermistors on selected gages were used to monitor temperature 

changes. Temperature readings were obtained at two locations on 

the steel channel beams at soldier beam No. 93, gage 1.3 and 3.1, 

as well as the first strain gage on each of the instrumented 

tiebacks. Reference is made to Graphs 17 and 18 which depict the 

temperature readings over the course of the monitoring period to 

date. 

As expected, the thermistors on the steel channel beams showed a 

direct response to ambient temperatures, once excavation 

progressed below the location of the gages. There is only minimum 

thermal separation between the ambient temperature and the 

location of the thermistor, the concrete surrounding the steel 

channel beams ,and the facing wall concrete. 

Temperature response was most clearly demonstrated by gage 3.1 

within the first 7-weeks of monitoring. This gage is located on 

the south flange at Section 3 of soldier beam No. 93. This 

location corresponds to the level at the bottom of wall footing 

elevation. The first two temperature readings at this location 

(67' and 60°F) reflected the cooling influence of the surrounding 

soil prior to excavation below the second level tieback location. 

Upon excavation to the bottom of wall footing elevation, 

temperature readings at the gage (3-weeks consecutively) were 

77'F. This represents a 17OF jump due to more direct gage 

exposure during excavation. This illustrates the effect of 

soldier beam exposure to ambient temperatures. Subsequent wall 

6-16 
footing construction and backfilling produced a 7' drop in 

temperature, again illustrating the thermal influence of the 

surrounding soil. 

The thermistors located at the approximate mid-point of the 

unbonded length of the tieback anchors showed relatively stable 

temperature readings. As would be expected, temperature 

fluctuations were much narrower then readings at the gages located 

on the soldier beam. Further, temperature readings at the tieback 

anchor locations, 58O to 68OF, yielded results expected for soil 

thermal conditions at considerable depth below surface grades. 

6-17 

?Qks = I I 

771 ! i 

I 

6-A-l 
lDP # JPILLcb. 5UA?T 

- 

PLAce CLAS5 ‘c’ 

CONCRlVt IhJ 

DRILLLO Sti~fT 

NO*: WALL F’AcINh UR 

hb3ER NOT 

SW FOR 

319 OR 34: b DRUB SHAFT 

CLARIN. 

TYPICAL WALL SECTlOhl 

%I= . I’-0’ 

Tieback Wall Details. 

6-A-2 
30’ OR 36-d’ ORILLZD SW 

Tieback Wall Details. 

6-A-3 
gOLD*CI 8L** ‘ TILOACR ScnrDuLq 

I 

-- 

-.- 

5 05b.5 03i.s *...I ;:i 

--- 

31.0 --_ 

: :::t: 

I I 31 __*_ 1 4.4 --- 

s1.5 .- - 5.0 e-v , 

8 

a‘. I 

S‘ 

10 Iluu 

34.4 ::: 

:: 670.3 

170.5 

34.46 5.1 1--- 

:: 870.) 

170.1. 

IS.2 15.0 ::: --- 

. . . . . 

35.3 5.4 --- 

--- 

--_ 

Tieback Wall Details. 

6-A-4 
,1 L ,,,1,,” ,/, “.I #,.Y 

1 ,E 

I” 

- 

- ; 

. 

.- 

. -. 

I 

1 _ 

I- 
CROSS SECTlOW 

RETA, ING UALL UNIT 71. S.E. 193 

NCffT It STREET GRADE SEPARATIGN 

C,TY OF LInA. OHIO 

ALLEN CWT’f 

*C.&c. 1/4”.1’-0” D1R. JAN.. 1988 FIGURE NO. 2 

,. . *. .-- 
INCLINOMETER Q 

SCALE: 3/4"=1'-0" 

30" DIAMETER S.B. 1193 

SECTION 1 

15 x 33.9 ELEVATION 868.0 

FACING WALL 

GAGE 2.3 GAGE 2.4 R 

YP.) V 

SCALE: 3/4"=1'-O" 

Cl5 x 33.9 S.B. t93 

CHANNELS SECTION 2 

ELEVATION 855.0 

CONCRETE 

-FACING WALL 

' 

GAGE 2.1 GAGE 2.2 

FIGURE 3A. STRAIN GAGE LOCATIONS 

SOLDIER BEAM #93 

6-A-7 
GAGE 3.3 

GAGE 3.5 

GAGE 3.4 

FLANGE (TYP.) 

w 

INCLINOMETER 0 

SCALE 3/4"=1'-0" 

GAGE 3.6 S..B- t93 

CJEB (TYP.) SECTION 3 

ELEVATION 850.0 

Cl5 x 33.9 (@ BOTTOM OF WALI 

FOOTING) 

3.1T 

GAGE 3.1 

30" DIAMETER 

DRILLED PIER 

GAGE 4.5 

INCLINOMETER 

ill 

SCALE 3/4"=1'-0" 

GAGE 4.6 

WEB (TYP.) S.B. 193 

SECTION 4 

ELEVATION 840.0 

Cl5 x 33.9 

DIAMETER 

DRILLED PIEP 

FIGURE 3B. STRAIN GAGE LOCATIONS 

SOLDIER BEAM #93 

6-A-8 
DOUBLF 

LOAD CELL 

DOUBLE CHANNELS 

SCALE: 1"=5' 

FIGURE 4. STRAIN GAGE & LOAD CELL LOCATIONS 

TYPICAL FOR INSTRUMENTED TIEBACKS 

I 
THE H. C. NUTTING COMPANY 

4120 AIRPORT ROAD 

CINCINNATI, OHIO 45226 

SOLDIER 

BEAM 

DATE LOCATION 

7-16-87 @ 24 & 93 

7-28-87 @ 24 6 93 

OHIO DEPARTMENT OF TRANSPORTATION 

DEMONSTRATION PROJECT NO.. 68, 

ALL-S.R. 81-16.83, NORTH STREET 

GRADE SEPARATION, 

LIMA, OHIO 

- Excavation level @ 1' below 1st tieback 

elevation (elev. - 867 @ 93, 866.5 @ 24) 

- Tiebacks not tensioned 

- Readings (background on SB 93 (beam gages) 

& 1st tieback @ 93 & 24, load cells @ 1st 

tieback level & inclinometers 

- Excavation level @ 2' below 2nd tieback 

elev. (Elev. 8.58 @ 24 & 93) 

- Tiebacks at 1st level tensioned 

Tieback elevation cd 24 - 867.5 @ 93 - 868.0 

- Readings (background) taken on 2nd level 

Tiebacks & load cells @ 24 & 93 prior to 

tensioning 

- Readings on remainder of gages also taken 

- Top of steel soldier beams 873.89 @ 93 

(Provided by ODOT) 873.90 @ 24 

- 2.9' top of beam :totop of clip @ 93, 2.7' 

@ 24, elev. 876.8 @ clip (93) 

elev. 876.6 @ clip (24) 

- A second reading was taken 7-28-87 

on all gages and load cells and inclinometer 

@ 93 after second tieback level tensioned 

- 2nd level tieback elev. 860 (24 & 93) 

- 2nd level tieback @ 24 not tensioned this date 

g-05-87 @ 24 & 93 - Excavation level @ 10'2 below 2nd tieback 

anchor levels 24 6 93 (Elev. 850.25) 

approximately same elev. as bottom of wall ftg. 

- All tiebacks tensioned - 2nd level tiebacks 

@ 24 tensioned subsequent to previous readings. 

- Readings taken all gages & load cells h 

inclinometers 

8-14-87 @ 24 & 93 

8-21-87 @ 24 & 93 

8-28-87 @ 24 & 93 

Same conditions as 8-05-87 

WaLl footing in place - bottom of footing 850.25 

@ 24, 850.0 @ 93 

- Facing wall @ 24 in place 

- Facing wall in at 93. 

- Subgrade for roadway complete adjacent to 

24 & 93, @ 24 - 854 elev. t, @ 93 - 853.5 elev. + 

9-4-87 @ 24 & 93 - Wall beam cap in place,top/cap 874.89 @ 93 

874.46 @ 24 

Figure 5. Log of field conditions on 

dates of instrumentation readings 
THE H. C. NUTTING COMPANY 

4120 AIRPORT ROAD 

CINCINNATI, OHIO 45226 

DATE LOCATION 

9-18-87 @ 24 & 93 

g-30-87 @ 24 & 93 

10-16-87 @ 24 & 93 

11-02-87 

12-08-87 

- Construction essentially completed 

- This & subsequent readings - routine 

- Routine Reading - 2 week interval 

- Routine Reading - begins 1 month reading intervals 

l-20-88 - Routine -ding 

2-22-88 -Routine Reading-begin threen-mthreading intervals 

5-27-88 - Routine Reading 

8-27-88 -Routine Reading 

SOLDIER 

BEAM 

OHIO DEPARTMENT OF TRANSPORATION 

DEMONSTRATION PROJECT NO. 68, 

ALL-S.R. 81-16.83, NORTH STREET 

GRADE SEPARATION, 

LIMA, OHIO 

- Road surfacing essentially complete 

Finish Grade @ 24 - 854.5 elev. 

@ 93 - 854.0 elev. 

- Storm sewer excavation made immediately 

behind wall @ 93, 2'? behind soldier beam 

pier, 4'? deep, bottom of trench @ elev. 870.5+ 

- Grade at back of wall @ 93 - 874.52, 

@ 24 - 874.0? 

Figure 5. (Continued) 

6-A-11 
THE H. C. NUTTINO CO. 

DIOITILT INCLINONETER SUUNARY SHEET 

NORTH STREET QRADE SEPARATION 

PROJECT NO. 9069.006 

INCLINONETER LOCATION NO. 24 HORIZONTAL SCALS: 1 INCH = .1 INCH 

TOP OF CASINO ELEV. 976.6 VERTICAL SCALE: 1 INCH = 10 FT. 

DATE OF REFERENCE READING: 7-16-97 

LATEST READING: READING NO. 11 DATE: 12-08-87 FILE: LINA25.11 

6-A-12 
THE H. C. NUTTINQ CO. 

DIOITILT INCLINOUETER SUNMARY SHEET 

NORTH STREET GRADE SEPARATION 

PROJECT NO. 9069.005 

INCLINOMETER LOCATION NO. 24 HORIZONTAL SCALE: 1 INCH = .l INCH 

TOP OF CASINQ ELEV. 879.8 VERTICAL SCALE: 1 INCH = lo FT. 

DATE OF REFERENCE READINO: 7-16-97 

LATEST READING: READINQ NO. 4 DATE: S-21-87 FILE: LINA25.4 

. 

,I 

1.7 

II 

I’ 

1. 

:: 

:: 

:: 

: 

, 

:I 1-l 

,* 17 IV 

** 

:: 

:: 2, 

:: 

3, 

6-A-13 
THE H’. C. NUTTINQ CO. 

DIDITILT INCLINONETER SUNNARY SHEET 

NORTH STREET ORADE SEPARATION 

PROJECT NO. 8069.005 

INCLINONETER LOCATION NO. 93 HORIZONTAL SCALE: 1 INCH = .I INCH 

TOP OF CASING ELEV. 876.8 VERTICAL SCALE: 1 INCH = 10 FT. 

DATE OF REFERENCE READING: 7-16-87 

LATEST READINO: READIND NO. 5 DATE: S-21-87 FILE: LINA93.5 

6-A-14 
THE H. C. NUTTINQ CO. 

DIQITILT INCLINOMETER SUMMARY SHEET 

NORTH STREET ORADE SEPARATION 

PROJECT NO. 9069.006 

INCLINOMETER LOCATION NO. 93 

TOP OF CASINO ELEV. 876.8 HORIZONTAL SCALE: 1 INCH = 91 INCH 

VERTICAL SCALE: 1 INCH = 10 FT. 

DATE OF REFERENCE READINQ: 7-16-87 

LATEST READINO: READING NO. 5 DATE: E-21-87 FILE: LINA93.5 

.--___----_-----_-__--------. 

DATI: 7-n-w 

DAIS: 1-*,-w 

DIW.: ,-m-11 

DATE: .-,.-.I 

DATE: 1-11-1, 

DEPTH 

IR., 

IIsT-“*‘I DIPLIRION. ,IYCWE‘, 

-., -.I -.I -.a -., 

:---------:---------;---------:---------;---------~---------~!-------~~ 

6-A-15 

70 

Kl 

50 

40 

30 

xl 

M 

cl 

0 

0 LC 

TIEBACK LOADS - S.B. #93 - IST LEVEL 

Y X 

I I I I I 

20 40 

+ 

TIME SINCE TIEBACK LOCKOFF [WEEKSI 

0 #2 A #3 X a4 

5 

2 

’ i? 

x 

w 
80 

70 

al 

50 

40 

3J 

20 

Kl 

cl 

TIEBACK LOADS - S.B. #93 - 2ND I .EVFeJ, 

LOAD TRANSFER WITH TlME 

3 A * 

h b c ” 

60 

cl LC 

TIME SINCE TEESACK LOCKOFF WEEKS 

+ #I 0 #2 A r3 X rt4 

GRAPH 2 

v * 
70 

60 

50 

40 

30 

20 

HI 

Cl 

TIEBACK LOADS - S.B. *24 - IST LEVEL 

LOAD TRANSFER WITH TIME 

h h 

h ” ” 0 

0 20 40 60 

TIMEi SINCE TIEBACK LOCKOFF WlXK!Sl 

0 LC + #I 0 #2 A #3 X r4 V -4s 

GRAPH 3 
60 

50 

40 

30 

20 

HI 

0 

-lo 

TIEBACK LOADS - S.B. #24 - 2ND LEVEL 

LOAD -I-RAN= WI-m -nME 

IEl 

I I I I I 

0 

cl LC 

GRAPH 4 

+ 

20 40 60 

#I 

-l-ME SINCE TIEBACK LOCKOFF 0 

0 tt2 A #3 X #4 
STRESS IN S.B. #93 CHANNELS - SECTION 1 

0.3 

0.1 

0 

-0.1 

-a3 

-0.4 I I I I I 

0 20 40 60 

lmlEslNcEtNmAL READINGS [WEEKS] 

0 N FIAN- CHANNEL, + NFLANGEAZCHANNEL 
1 

a9 

a8 

a7 

a6 

a4 

a3 

aI 

0 

-0.1 

STRESS IN S.B.#93 CHANNELS - SECT.2,3,4 

I 

I 

7 

h A h 

0 20 40 60 

TlME SINCE tNlTIAL FEADNGS IwEEKs] 

cl SECT.2 + SECT.3 0 SECT.4 

GRAPH 6 
1 

a9 

as 

0.7 

Cl.6 

0.5 

a4 

a3 

a2 

al 

cl 

xl.1 

STRESS IN S-B.*93 CHANNELS - SECT.2,3,4 

q 

0 20 40 60 

0 sEcT.2 

T-ME SINCE INITIAL FEAIMNGS 0 

+ SECT.3 0 SECT.4 
60 

so 

40 

30 

20 

HI 

0 

-Jo 

LOADS IN COMPOSITE S.B.#93 -SECT.2,3,4 

0 20 40 60 

TME StNCE lNTTlAL READINGS [WEEKS] 

El SECT.2 + SECT.3 V SECT.4 

GRAPH 8 
AXIAL LOAD ‘IRANSFERWl-IHTlME-32GC 

LOADS IN S.B.#93 CHANNELS - SECT.2,3,4 

17 - 

16- 

15 - 

18 , 

14 - 

13 - 

anKl9- 

8- 

7- 

6 

5 

4 

3 

2 

1 

0 

-1 

0 20 40 60 

TIME SINCE INITIAL READINGS WEEKS 

cl sEcr2 + ECT.3 0 SECT.4 
LOADS IN S.B.#93 CHANNELS - SECT.2,3,4 

AXAL LOAD TRANsFERwlmTlME 

18 

17 

16 

15 

14 

13 

a 

11 

la 

9 

8 

7 

I I I I I 

0 20 40 60 

TlME SINCE INlTlAL READDiGS WEEKSI 

Cl SECT.2 + SECT.3 0 SJxT.4 

GRAPH 10 
0 

d 

cl 

1 

Edlw) ‘BTS XLLISOdMO3 NI tlV0-l -IVKV 

~-B-II 
6 

5 

4 

3 

2 

1 

0 

-1 

-2 

MOMENTS IN S.B.#93 CHANNELS - SECT. 2 

0 20 40 60 

El Y-AXIS 

TIME SINCE INITIAL REAXXNGS [wEEKsI 

+ X-AXE/W 0 X-AXE/E 

GRAPH 12 
KdDl-JAI STINNVH3'WSNI S.LNJl4Ol4 

6-0-13 





REPORT NO. 5 

PERMANENT GROUND ANCHORS FOR LATERAL 

SUPPORT OF BRIDGE END SLOPES 

DEMONSTRATION SITE 

DIMOND BOULEVARD UNDERPASS STRUCTURE 

PROJECT NO. FM-0520(l) 

SUBMITTED BY: 

ALASKA DEPARTMENT OF TRANSPORTATION 

TO THE FHWA DEMONSTRATION PROJECTS DIVISION 

FOR INCLUSION IN 

FHWA DEMONSTRATION PROJECT NO. 68 

PERMANENT GROUND ANCHORS 

ANCHORAGE, ALASKA 

NOVEMBER 1986 

TABLE OF CONTENTS 

I. INTRODUCTION ............................................ 5-l 

II. RETAINING WALL DESIGN ................................... 5-l 

III. RETAINING WALL CONSTRUCTION ............................. 5-2 

IV. RETAINING WALL INSTRUMENTATION .......................... 5-4 

V. DATA ANALYSIS ........................................... 5-5 

VI. ADDITIONAL WORK ......................................... 5-6 

5-i 

I. INTRODUCTION 

A permanent ground-anchored retaining wall has been constructed on 

the Dimond Boulevard Project. The widening of Dimond Boulevard 

beneath the new Seward Highway overpass structure required the 

removal of a portion of the existing bridge end fill slope. This 

slope supported the abutments 

footings. 

which were founded on spread 

The only retaining wall that was possible to construct 

without a detrimental effect to the overpass structure was a 

soldier pile wall using permanent soil anchors. To our knowledge, 

this was only the second time that permanent ground anchors were 

used for a retaining wall supporting a bridge structure founded on 

spread footings in the United States. The other case was in 

Washington State. 

The primary purpose of this project is to: 

1. Evaluate the performance of permanent ground-anchored 

retaining walls supporting and existing bridge structure 

founded on spread footings. 

2. Measure 

after the 

the bridge settlement that takes place during and 

installation of the retaining wall. 

3. Provide data for incorporation into the Federal 

Administration's 

Highway 

Permanent Ground Anchor Demonstration 

Project No. 68. 

This progress report summarizes the installation of the 

instrumentation, initial 

during the 

data analysis, and bridge performance 

first year after construction. The final 

scheduled for January 

report 

1989 will present all results and 

conclusions of the test program. 

II. RETAINING WALL DESIGN 

The New Seward Highway overpass was originally constructed in 

1976. This structure is a 110 foot long, single span, prestressed 

concrete girder structure supported on spread footings founded in 

approach fills. In 1986, Dimond Boulevard was widened from 66 

feet to approximately 96 feet. To increase the headroom 

clearance, the Dimond Boulevard grade was also lowered two feet, 

which required an eight foot subcut below the 

remove pockets of buried peat. 

existing grade to 

The grade and template changes to Dimond Boulevard required the 

construction of two permanent retaining walls. 

retaining wall supported each abutment. 

A 2124 square foot 

permanent wall is 12 feet; however, 

The maximum height of the 

the wall was constructed to a 

5-1 
height of 18 feet to support the walls of the sub-excavation. 

This increased the total wall area to approximately 

feet for each wall. 

3100 square 

The retaining wall was designed by the Alaska Department of 

Transportation and Public Facilities. A performance specification 

was used for the permanent ground anchors. Based on this 

specification, the contractor was required to: 

1. meet an experience-based pre-qualification requirement, 

2. design the ground anchors including method of installation 

and bond length, 

3. proof 

and 

test each anchor to 150 percent of the design load, 

4. conduct performance tests on the first four anchors to 150 

percent of the design load. 

The soil conditions at the location of the soldier piles were 

primarily dense sandy gravel to gravelly silty sand with 

occasional cobbles 

sand (glacial till). 

(backfill material) overlying gravelly silty 

had 

This area was previously a peat bog which 

been excavated during the initial construction of the 

overpass. Pockets of peat, which were located at the edge of the 

existing structure, were also removed as part of this project. 

The water table was 7 feet below the ground surface. 

The soils in the vicinity of the bond zone of the ground anchors 

were primarily very dense sandy gravel to gravelly sand (backfill) 

overlying silty sandy gravel (glacial till). 

III. RETAINING WALL CONSTRUCTION 

The HP 12 x 84 soldier piles were installed through holes drilled 

in the bridge deck. 

6 inches, and 

The soldier piles were spaced at 7 feet, 

hammer instead of 

embedded below ground using an ICS 812 vibratory 

pre-boring the holes. 

feet below original ground, 

The piles were driven 30 

or 22 feet below the bottom of the 

subcut. A template consisted of a 12 inch tubular steel frame 

supporting angle clips which restrained the pile flanges was used 

to maintain alignment. 

Alignment of the driven piles was true within l/2 inch. Some 

difficulty in obtaining exact alignment and elevation required 

that piles be occasionally extracted part way and re-driven. This 

was easily accomplished with the vibratory hammer. The only 

problem encountered was friction heating of the angle clips due to 

passage of the pile flange. The vibration peak particle velocity 

5-2 
measured at the footings ranged from .06 to .12 inch per second. 

The ground anchors were inclined at 15 degrees below horizontal. 

The free (no bond) zone was 20 feet long and the bond zone was 

25.5 feet long (Figure 1). The design load of the anchors was 105 

kips. The anchors had single level corrosion protection, i.e. the 

grout was the only corrosion protection. The anchor rod was 

protected from corrosion in the bond zone by grout encapsulation. 

The free zone was protected by a grease-filled PVC sheath. 

The ground anchors were installed with a rotary air percussion 

drill. In most cases, the 5-l/2 inch diameter casing was driven 

with a conical tip. The casing was then disengaged from the tip 

and extracted by a hydraulic jack during the grouting operation. 

In some cases, the casing was driven open-ended and drilled out. 

Dywidag thread bars, l-3/8 inch in diameter, were placed in the 

casing. (The contractor had the option of using steel strands or 

steel bars.) No coupler connections were needed for the bars. A 

l/2 inch PVC sheath tubing was attached to the bars for secondary 

grouting. Two holes drilled in the tubing and covered with a 

rubber compound performed as valves in the bond zone. A 20-foot 

long PVC sheath was placed around the bar in the free zone to 

prevent bonding between the grout and the bar. 

The grout was mixed at a ratio of 7 gallons of water per bag of 

grout. The grout was pumped through the grout tube at 100 psi 

pressure as the casing was extracted. After 24 hours, an attempt 

was made to inject additional grout at 500 to 600 psi. Only 39 

percent of the anchors accepted secondary grouting. The volume 

for secondary grouting ranged from 10 to 50 gallons. Some of the 

grout tubes were damaged during installation or blocked by 

concrete shards inadvertently allowed into the tube. 

The first four anchors were performance tested to 150 percent of 

the design load. The remaining anchors were proof tested to 150 

percent of the design load. All of the anchors were tested from 3 

to 14 days after primary grouting. All anchors passed their test 

on the first try. Due to problems with the load cells, the test 

loads and the lockoff loads were based on the "calibrated" jack 

readings. 

The timber lagging was a dense Douglas Fir which was pressure 

treated with creosote. The lagging was cut in the field to fit 

between the piles. The ends were field treated with creosote. 

The excavation was not allowed to proceed beyond 18 inches below 

the bottom of previously placed lagging before placing a new lag. 

In late August, a severe rainstorm struck the project. Water 

funneled through the deck holes, washed material from behind the 

retaining wall. This resulted in 1 to 2-l/2 inches of settlement 

in the overpass footings. The settlement occurred when the 

5-3 
excavation was only 6 feet below the footing prior to the 

installation of the ground anchors. To prevent additional 

settlement, grout was pumped behind the walls to fill the voids. 

This grouting operation continued as the excavation and wall 

installation proceeded. 

Precast concrete panels were attached by four hangers welded to 

the soldier piles. The panels were difficult to install due to 

the small hangers, uneven deflection 

loads and 

in piles from the anchor 

panels. 

insufficient head room to use a boom to lift the heavy 

The project personnel suggested that an adjustable double 

angle hanger would ease the installation. 

The average total cost of the retaining wall without the precast 

concrete wall panels was approximately $69 per square foot. The 

precast panels for the exposed portion of the retaining wall cost 

an additional $18 per square foot installed. 

The contract prices were: 

Item Quantitv 

Wall Anchors 38 ea. 

Deck Holes 26 ea. 

Precast Wall Panels 4,248 sf. 

Aesthetic Facia 4,248 sf. 

Steel Piles 1,944 If. 

Treated Timber 48.74 MBM 

Unit Price 

$2,500 $ 95,000 

2,000 52,000 

15 63,270 

3 12,744 

100 194,400 

1,500 73,110 

TOTAL $490,524 

IV. RETAINING WA&L INSTRUMENTATION 

A SINCO Model 513510 electric load cell was installed at the head 

of one anchor on each wall. A SINCO Model 51309 electric load 

cell readout device was purchased to monitor the load cells. 

Vibrating wire strain gauges (SINCO Model 52621 with pickup sensor 

Model 52622) were installed to measure load transfer along the 

anchor shaft (bond zone) and to measure strain in the no load 

zone. Three sets of 

anchors, 

two gauges each were installed on the two 

feet and 

one set near the anchor head; two sets approximately 4.9 

13.2 feet from the ,end of the bond zone. A vibrating 

wire strain indicator readout divide (SINCO Model 52669) was used 

to monitor the strain gauges. Of the 12 strain gauges installed 

only the two strain gauges located in the middle of the bond zone 

on Anchor #16 failed to work. 

5-4 
The proposed slope inclinometer casing could not be installed 

since the soldier piles were driven and not 

holes. 

placed in pre-bored 

v. DATA ANALYSIS 

Based on the data 

installation, 

collected during the first year after 

the following observations can be made: 

1. There is a lo-15 kip difference in the measured load 

between the load cell and the strain gauges in the no bond 

zone on Anchor #16. This may be due to either 

miscalibration of the strain gauges, bearing plates 

adjacent to the load cell may be too thin to transfer load 

without bending, or the grout has extended into the no 

bond zone which permits some load transfer to the no bond 

zone. 

2. Anchor #35 was originally tensioned to 105 kips on g/11/85 

and de-tensioned on the same day. The anchor was 

tensioned again to 105 kips on g/12/86; however, six days 

later the anchor head was damaged. The anchor was 

repaired and re-tensioned a third time. The initial 

strain gauge readings prior to anchor stressing were 

different each time. The lowest set (second) of readings 

were used as initial readings for this report. 

3. The load cell and strain gauges in the unbonded zone on 

Anchor #16 indicate an apparent load reduction rate of 5 

to 6 kips per log cycle of time from 

days to 300 days. 

the period to 100 

This equates to a load reduction of 10 

to 12 kips in thirty years. The load cell and strain 

gauges in the unbonded zone of Anchor #35 indicate an 

apparent load reduction of 1.25 to 1.4 kips per log 

time. 

cycle 

The load reduction may be caused by movement of the wall 

into the backslope, due to the applied anchor load 

exceeding the at-rest pressure behind the wall. If this 

is the case, the load reduction should cease when the 

anchor load equals the at-rest pressure. 

4. The difference 

and 

in anchor load measured by the load cells 

strain gauges with the load measured by the 

"calibrated" jack may be due to inaccurate load 

measurement of the anchor loads by the jack. 

5. The strain gauges located 4.9 feet into the bond zone 

indicate a load reduction while the strain gauges located 

13.2 feet into the bond zone indicates an increase in 

load. This is due to load transfer down the bond zone 

over time. 

5-5 
The following tasks will be performed during the next two years: 

1. Continue monitoring the load cells and strain gauges for 

the next two years. 

2. Retension the two instrumented ground anchors to determine 

the cause for the difference in load measured between the 

strain gauges in the no bond zone and the load cell. 

3. Establish and monitor survey control points on the soldier 

piles to determine lateral movement. 

4. Reestablish and monitor survey control points on the 

overpass structure to measure settlement. 

5. Prepare final report to include data interpretation, 

conclusions, and recommendations. 

5-6