National Academies Press: OpenBook
« Previous: Chapter 1 - Introduction and Research Approach
Page 11
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 11
Page 12
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 12
Page 13
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 13
Page 14
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 14
Page 15
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 15
Page 16
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 16
Page 17
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 17
Page 18
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 18
Page 19
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 19
Page 20
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 20
Page 21
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 21
Page 22
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 22
Page 23
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 23
Page 24
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 24
Page 25
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 25
Page 26
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 26
Page 27
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 27
Page 28
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 28
Page 29
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 29
Page 30
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 30
Page 31
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 31
Page 32
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 32
Page 33
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 33
Page 34
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 34
Page 35
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 35
Page 36
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 36
Page 37
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 37
Page 38
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 38
Page 39
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 39
Page 40
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 40
Page 41
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 41
Page 42
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 42
Page 43
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 43
Page 44
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 44
Page 45
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 45
Page 46
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 46
Page 47
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 47
Page 48
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 48
Page 49
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 49
Page 50
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 50
Page 51
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 51
Page 52
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 52
Page 53
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 53
Page 54
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 54
Page 55
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 55
Page 56
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 56
Page 57
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 57
Page 58
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 58
Page 59
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 59
Page 60
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 60
Page 61
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 61
Page 62
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 62
Page 63
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 63
Page 64
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 64
Page 65
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 65
Page 66
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 66
Page 67
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 67
Page 68
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 68
Page 69
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 69
Page 70
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 70
Page 71
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 71
Page 72
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 72
Page 73
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 73
Page 74
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 74
Page 75
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 75
Page 76
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 76
Page 77
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 77
Page 78
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 78
Page 79
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 79
Page 80
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 80
Page 81
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 81
Page 82
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 82
Page 83
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 83
Page 84
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 84
Page 85
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 85
Page 86
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 86
Page 87
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 87
Page 88
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 88
Page 89
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 89
Page 90
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 90
Page 91
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 91
Page 92
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 92
Page 93
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 93
Page 94
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 94
Page 95
Suggested Citation:"Chapter 2 - Findings." National Academies of Sciences, Engineering, and Medicine. 2006. Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing. Washington, DC: The National Academies Press. doi: 10.17226/13936.
×
Page 95

Below is the uncorrected machine-read text of this chapter, intended to provide our own search engines and external engines with highly rich, chapter-representative searchable text of each book. Because it is UNCORRECTED material, please consider the following text as a useful but insufficient proxy for the authoritative book pages.

11 CHAPTER 2 FINDINGS FINDINGS FROM LITERATURE STUDY Over the past two decades, GRS has been used in the con- struction of various earth structures, including retaining walls, embankments, slopes, and shallow foundations. In actual construction, GRS structures have demonstrated many distinct advantages over their conventional counterparts. GRS structures typically are more ductile, more flexible (hence more tolerant to differential settlement), more adapt- able to low quality backfill, easier to construct, and more economical. They also require less overexcavation. In recent years, applications of the GRS technology to bridge-supporting structures have gained increasing atten- tion. The facing of GRS bridge-supporting structures can be grouped into two types: rigid and flexible. A rigid facing is a continuous reinforced concrete facing, either precast or cast- in-place. A flexible facing, on the other hand, typically takes the form of wrapped geosynthetic sheets, dry-stacked con- crete modular blocks, natural rocks, or gabions. In contrast to a flexible facing, a rigid facing offers a certain degree of global bending resistance along the entire height of the fac- ing, thus offering greater constraint to lateral earth pressure- induced “global” bending deformation. Since 1994, the Japan Railway has constructed many full- height facing GRS bridge abutments and piers (e.g., Tateyama et al., 1994; Kanazawa et al., 1994; Tatsuoka et al., 1997) using a rigid facing GRS wall system developed by Tatsuoka and his associates at the University of Tokyo. These GRS bridge- supporting structures have been constructed in two stages. The first stage involves constructing a wrapped-faced GRS wall with the aid of gabions, and the second stage involves casting in place a full-height reinforced concrete facing over the wrapped face. Field measurement has shown that these structures experienced little deformation under service loads and have performed far better than conventional reinforced concrete retaining walls and abutments in the 1995 Japan Great Hansin earthquake that mea- sured 7.2 on the Richter scale (Tatsuoka et al., 1997). Most recently, Tatsuoka and his associates developed a preload- prestress method for improved performance of the GRS bridge- supporting structures (Tatsuoka et al., 1997; Uchimura et al., 1998). Despite their success, the rigid facing GRS bridge- supporting structures have found applications only in Japan, mostly because of their higher cost and longer construction time compared with GRS walls with flexible facings. GRS bridge-supporting structures with flexible facings have been the subject of many studies and recently have seen some actual applications, in the United States and abroad. This study synthesizes the measured behavior and experiences gained from case histories of flexible facing GRS bridge- supporting structures from around the world. Observations were made in relation to performance, design, and construc- tion of flexible facing GRS bridge-supporting structures. The case histories were organized into two groups: in-service struc- tures and field experiments. Most of these studies were on bridge abutments, with a few on bridge piers. The design and construction of GRS bridge abutments are similar in principle to GRS walls, except the former typically are subject to a rather high surface load close to the wall face. Also, some U.S. states do not permit the use of segmental concrete facing in GRS bridge-supporting structures because of concerns with the durability of masonry units when exposed to chemical agents such as de-icing fluids. Based on the measured perfor- mance of the case histories, observations were made in rela- tion to performance, design, and construction of GRS bridge- supporting structures. Some of the material properties and the methods for determining the properties are not reported because they are not available in the source materials. In-Service Bridge-Supporting Structures The construction-related information and measured perfor- mance of six in-service GRS bridge abutments are described below. The six abutments are the Vienna railroad embank- ment in Austria (Mannsbart and Kropik, 1996), the New South Wales GRS bridge abutments in Australia (Won et al., 1996), the Black Hawk bridge abutments in Colorado (Wu et al., 2001), the Founders/Meadows bridge abutments in Col- orado (Abu-Hejleh et al., 2000), the Feather Falls Trail bridge abutments in California (Keller and Devin, 2003), and the Alaska bridge abutments in Alaska (Keller and Devin, 2003). Case A1: Vienna Railroad Embankment, Austria (Mannsbart and Kropik, 1996) A temporary GRS embankment was constructed in Vienna, Austria, to support a railroad track. The railroad embankment had a height of 2.1 m and a slope inclination of

63 deg from the horizontal. A needle-punched nonwoven geotextile was used as the reinforcement. The geotextile had a tensile strength of 23 kN/m with elongation at break of 45 percent. The reinforcement spacing and length were 0.3 m and 1.7 m, respectively. The backfill was a compacted grav- elly sand. Its placement unit weight was 21 kN/m3, and the design internal friction angle was 35 deg. The individual layers of the structure were built using a removable formwork consisting of steel angles and wooden bars. To get adequate friction between the adjacent geotextile layers, a thin layer of sandy gravel was placed on each lift before the installation of the next layer. Given that the struc- ture had to fulfill only a temporary function, a wrapped-around wall face was used and the surface protection was omitted. Above the reinforced structure, a 0.9-m-high unreinforced embankment with a slope of 45 deg was built as a buffer for the traffic. The design traffic load was 60 kPa, exerted at 1.45 m from the top edge of the unreinforced embankment. The cross-section of the temporary embankment is shown in Figure 2-1. Weekly settlement measurement was carried out on 6 points along the 100-m-long embankment. The results indicated that under traffic load, the measured settlement was nil at four of the six points, and at the other two points the settlement was less than 1 mm. Case A2: New South Wales GRS Bridge Abutments, Australia (Won et al., 1996) Geogrid reinforced bridge abutments with a segmental block facing were constructed to support end spans directly for a major bridge in New South Wales, Australia, in 1994. The bridge consisted of a nine-span superstructure over the Tweed River. The abutments were up to 10 m high, con- structed in a terraced arrangement, as shown in Figure 2-2. The facing comprised “Keystone” segmental concrete blocks that were partially voided internally, and aggregates were used to fill the block during construction. High-strength 12 fiberglass dowels were used to interlock block layers verti- cally. Foundation conditions at the site consisted of a 1- to 3-m-thick layer of loose silty sand containing thin discontin- uous silty clay layers overlying a medium dense silty sand layer varying in thickness from 7 m to 10 m. Sandstone bedrock was present at 13 m depth. The two abutments are referred to as Abutment A and Abut- ment B. Abutment A consisted of three terraced segmental block walls with 12 layers of a Tensar HDPE geogrid, SR 110, beneath the sill beam. The tensile strength of the geogrid was 110 kN/m at 11.2 percent strain. Total tiered height was 6.5 m. Abutment B consisted of four terraced segmental block walls with 17 layers of SR110 geogrid beneath the sill beam. Total tiered height was 9.5 m. To account for creep, temperature vari- ation, and construction damage, the allowable long-term design strength for the SR110 geogrid was taken as 27 kN/m. The ver- tical spacing of geogrid layers was 40 cm or 60 cm. The maxi- mum reinforcement length was 15 m. The backfill material, a fine sand, was compacted to at least 95 percent Standard Rela- tive Density to have a design friction angle of 32 deg. Addi- tional layers of geogrid, 5 m long with a wrap-around face, were used to reduce active earth pressure behind the sill beam. The unreinforced concrete sill beam was 20 cm thick and 2.5 m wide. It was set back 2.5 m from the edge of the top wall to reduce the effects of horizontal pressure because of sill beam load distribution through the reinforced soil. In view of the loose nature of the foundation soil, the top 1 m was excavated and compacted in the vicinity of the lowest-tiered wall. A comprehensive monitoring program was implemented to evaluate the performance of Abutment B. Sill beam loading occurred during January 1994. The maximum reinforcement tension at Level 1 approached 33 kN/m and occurred toward the back of the reinforced soil block. The maximum rein- forcement tension at Level 2 was 21 kN/m and occurred toward the back of the reinforced soil block. At Level 3 rein- forcement, the effect of sill beam loading was evident with a maximum reinforcement tension of 22 kN/m occurring under the sill beam region. The maximum strain in the geogrid was 1.6 percent, occurring at Level 1. The maximum settlement was 80 mm. Lateral movements of the reinforced soil structure deduced from wall survey and inclinometers I1 and I2 (see Figure 2-2) were 10 mm up to the completion of the abutment and 26 mm post construction movements for the lowest-tiered wall. Subsequent site investigations of the loose upper silty sand layer indicated the presence of thin discontinuous seams of medium stiff silty clay, which could have contributed to the deformation response at the base of the structure. Case A3: Black Hawk Bridge Abutments, Colorado (Wu et al., 2001) Two rock-faced GRS abutments were constructed to sup- port the Bobtail Road Bridge, a 36-m-span steel arched bridge in Black Hawk, Colorado (see Figures 2-3 and 2-4). Figure 2-1. Cross-section of the Vienna railroad embankment, Austria (Mannsbart and Kropik, 1996). 0.9 m 2.1 m 6% 1.7 m 63° 0.3 m 0.9 m 1.45 m Sleeper 2.6 m p = 60 kN/m2

13 Compacted Sand Fill 2.5 m Bridge Sill Beam Inclinometer I2 "Keystone" Blocks 7 m 2 m 2 m Inclinometer I1, I2 Vertical Borehole 2 m 2 m Inclinometer I1 Tensar SR110 15 m 13 m 0.4 m Spacing Geogrid Level 1 Level 2 Tensar SR80 5 m 11 m 11 m 5 m Level 3 0.6 m Spacing Geogrid Figure 2-2. Cross-section of the New South Wales GRS bridge abutments, Australia (Won et al., 1996). Figure 2-3. Cross-section of the Black Hawk bridge abutments (Wu et al., 2001).

Each GRS abutment comprised a two-tier GRS mass with two square footings on the lower tier and a strip footing on the upper tier. The square footings on the West abutment are referred to as Footings #1 and #4, and the square footings on the East abutment are referred to as Footings #2 and #3. The GRS bridge abutments were constructed on a stiff soil. The thicknesses of the lower tier reinforced soil mass under Footings #1 and #4 were, respectively, 4.5 m and 1.5 m; and 7.5 m and 1.5 m under Footings #2 and #3, respectively. The lower part of the GRS abutment was embedded in the ground, while the upper part was above ground. Only the part above ground was constructed with rock facing. The above 14 ground portion of the abutment had different heights, vary- ing from 1.0 m to 2.7 m for the West abutment; and from 1.0 m to 5.4 m for the East abutment. The thickness of the upper tier reinforced soil mass was 1.8 m for both abutments. The upper tier reinforced soil mass was built to support the strip footing and the approach ramp. The abutments were constructed with the onsite soil, clas- sified as SM-SC per ASTM D2487, and reinforced with lay- ers of a woven geotextile at vertical spacing of 0.3 m. The polypropylene woven geotextile (Amoco 2044) had a wide- width tensile strength of 70 kN/m in both machine and cross- machine directions at 18 percent strain, per ASTM D4595 Figure 2-4. Footings and foundations of the Black Hawk bridge abutments (Wu et al., 2001).

(the wide-width strip method). The backfill had 12 percent of fines (passing sieve No. 200). The backfill material was compacted to 91 percent relative compaction per AASHTO T-99 (the moisture-density relation of soil was determined by using a 2.5 kg rammer with a 305 mm drop), having a dry unit weight of 15.8 kN/m3 at a water content of 12.2 percent. The measured friction angle and cohesion, as determined from the CD triaxial compression tests, were 31 deg and 34 kPa, respectively. For each square footing, a vertical pressure of 245 kPa (1.6 times the design load of 150 kPa) was applied and sus- tained for 100 minutes, then unloaded to zero. Three loading- unloading cycles were applied following the first loading- unloading cycle. In the reloading cycles, the typical applied pressure was the design load (150 kPa). For the strip footing, the vertical load was increased incrementally to 80 kPa (2 times the design load of 40 kPa), sustained for 120 minutes, and then unloaded to zero. The vertical load applied in the reloading cycle was 40 kPa (the design load). The load was maintained for 120 minutes before unloading. At the design load of 150 kPa in the preloading cycle, the average settle- ments were 13.3 mm, 6.4 mm, 28 mm, and 4.9 mm for Foot- ings #1 through #4, respectively. At 150 kPa in the first reloading cycle, the average settlements were reduced to 2.5 mm, 3.8 mm, 4.5 mm, and 3.3 mm for Footings #1 through #4 respectively. Further reduction in the settlement was neg- ligible in the subsequent reloading cycle. Preloading reduced the maximum lateral movement at 150 kPa loading pressure from 1.5 mm to 0.6 mm in Footing #1, and from 13.2 mm to 4.5 mm in Footing #3. In the preloading cycle, under a load of 245 kPa sustained for 60 minutes, the vertical creep displacements of Footings #1 to #4 were, respectively, 6.7 mm, 4.0 mm, 7.2 mm, and 2.1 mm. In the reloading cycle, under the sustained load of 150 kPa, the vertical and lateral creep deformations were insignificant. At 80 kPa in the preloading cycle, the maximum strains in layers A, B, and C were 0.18 percent, 0.04 percent, and 0.06 percent, respectively. At a sustained load of 80 kPa in the preloading cycle, the creep strains in layers A, B, and C were 0.032 percent, 0.009 percent, and 0.003 percent, respec- tively. Locations of layers A, B, and C are shown in Figure 2-3. The creep strains were negligible at the sustained load of 40 kPa in the reloading cycle. Based on the measured data, the following findings and conclusions were made: • By preloading the reinforced soil mass to 245kPa, the settlement at the design load of 150 kPa was reduced by a factor of 1.5 to 6 for the four square footings. • Preloading also reduced the lateral movement of the GRS abutments. The lateral movement was reduced by a factor of 2.5 to 3 at 150 kPa. • After the first reloading cycles, there was no significant reduction of lateral and vertical displacements of GRS abutments in the subsequent reloading cycles. 15 • The maximum strain mobilized in the reinforcement was very small (less than 0.2 percent at 80 kPa). • Preloading reduced creep strains in the reinforced struc- ture and the geotextile reinforcement. Case A4: Founders/Meadows Bridge Abutments, Colorado (Abu-Hejleh et al., 2000) A replacement bridge was constructed over Interstate Highway 25 at Founders/Meadows Parkways near Castle Rock, Colorado, in 1999. In this bridge abutment, both the bridge and the approaching roadway were supported by a system of GRS segmental retaining walls. The front GRS wall supports the bridge superstructure, which extents around a 90-deg curve into a lower GRS wall supporting the wing wall and a second tier, the upper GRS wall. The GRS abutment was constructed on the native claystone or sand- stone bedrock. The plan view of the structure is shown in Fig- ure 2-5. Each span of the bridge was 34.5 m long and 34.5 m wide. The design of the abutment followed the AASHTO (1997) guidelines. Figure 2-6 shows the typical cross-section of the abutment. For the reinforced soil zone behind and below the bridge abut- ment, a trapezoid-shaped reinforcement was adopted, in which reinforcement increased linearly from 8.0 m at the bot- tom with 1H:1V slope toward the top. The reinforcement length for the abutment wall was 11 m to 13 m. The center- line of the bridge abutment wall and edge of the foundation were 3.1 m and 1.35 m from the front of the wall face. Dry- stacked hollow-cored concrete blocks were used as the fac- ing. The lower wall had a maximum height of 4.5 m to 5.9 m and the upper wall had a maximum height of 3.0 m for the West abutment and 3.2 m for the East abutment. The lower wall had a minimum embedment of 0.45 m. The abutment was constructed in two phases to accommodate traffic needs. Three grades of geogrid reinforcement were used: UX6 with an ultimate strength of 157.3 kN/m used below the foun- dation, UX3 and UX2 with ultimate strengths of 64.2 kN/m and 39.3 kN/m, respectively, per ASTM D4595, used behind the abutment wall. The ultimate strength of the geogrids was measured in accordance with the ASTM D4595 test method. The reinforcement spacing was 0.4 m. The backfill soil was a mixture of gravel (35 percent), sand (54 percent) and fines (11 percent). The average unit weight and dry unit weight of the compacted fill were 22.1 kN/m3 and 21 kN/m3 (95 per- cent of AASHTO T-180, the moisture-density relation being determined by using a 4.54 kg rammer with a 457 mm drop), respectively. The average placement moisture content was 5.6 percent. Field monitoring was performed with various instruments during and after the construction of the structure. The mea- sured vertical stresses did not differ significantly from the static states calculated as σz = γz + q + Δσz, where q is the uniform surcharge and Δσz is the increase in vertical stress caused by concentrated surcharge loads assuming 2V:1H

pressure distribution. The horizontal stresses measured on the facing at the end of construction, however, were much smaller than the Rankine active earth pressures. The mea- sured geogrid strains at the end of construction were very low, on the order of 0.1 percent. The measured outward movement of the GRS wall face was also very small. The maximum outward movement experienced along Section 400 during the construction of the front GRS wall up to the bridge foundation elevation was about 9 mm. The maximum outward movements experi- enced during placement of the bridge superstructure were on the order of 7 mm to 9 mm. The field measurements also indicated the sill settled about 13 mm because of the loads of the bridge and the approaching roadway structures. Along Section 400 (see Figure 2-5), the leveling pad settled verti- cally almost 5 mm during the construction of the front GRS wall up to the bridge foundation elevation and settled another 6 mm when the bridge and approaching roadway structures were placed. Post-construction performance of the Founders/Meadows bridge abutment was evaluated by Abu-Hejleh et al. (2002), with the following findings: 16 • Eighteen months after opening to traffic, the maximum outward displacement of the front wall facing and the maximum settlement of the bridge abutment footing were 13 mm (0.22 percent of wall height) and 11 mm (0.18 percent of wall height), respectively. The maxi- mum outward displacement of the front wall facing occurred at the elevation directly below the bridge sill. • Movement of the leveling pad (located at the base of the GRS structure) was negligible, and the outward wall dis- placement tended to decrease toward the leveling pad. • Both the rates of wall movements and the strain of geogrid reinforcements decreased with time. • Outward wall displacement as inferred by the integra- tion of the strain distribution curve with respect to the reinforcement length matched closely with that deter- mined from surveying. This implies that little slippage between the soil and reinforcement had occurred. • Probable causes for post-construction movements were traffic load, deformation under sustained load (creep), and seasonal variation. • The GRS bridge abutment shows no sign of the “bridge bump” problem. The Founders/Meadows GRS bridge Figure 2-5. Plan view of the Founders/Meadows bridge-supporting structure (Abu-Hejleh et al., 2000).

abutment has exhibited excellent short- and long-term performance characteristics. Abu-Hejleh et al. (2003) also indicated that the rate of creep reinforcement strain under service load decreased with time: a maximum increase of strain of 0.09 percent during the first year, a maximum increase of strain of 0.04 percent dur- ing the second year, and a maximum increase of strain of 0.02 percent during the third year in service. The largest rein- forcement strain occurred directly beneath the bridge sill. The maximum reinforcement strain after about 33 months in service was 0.27 percent. The Colorado DOT concluded that the general layout and design of future GRS abutments should follow those in the Founders/Meadows abutment. The GRS abutments work well for multiple span bridges, have the potential for elimi- nating the “bump at the bridge” problem, avoid disadvan- tages associated with the use of deep foundations, and allow for construction in stages and within a smaller working area. 17 The Colorado DOT provided the following guidelines for design and construction of GRS abutments: 1. The foundation soil for these abutments should be firm enough to limit the post-construction settlement of the bridge sill to 75 mm. 2. The designer should plan for a bridge sill settlement of at least 25 mm caused by the bridge superstructure loads. 3. The maximum tension line needed in the internal sta- bility analysis should be assumed bilinear, starting at the toe of the wall and extending through a straight line to the back edge of the bridge sill at the mid height of the wall, and from there extending vertically to the back edge of the bridge sill. 4. Ideally, construction should take place during the warm and dry seasons. 5. The backfill behind the abutment wall should be placed before the girders. Figure 2-6. Typical cross-section of front and abutment walls, the Founders/Meadows bridge abutments (Abu-Hejleh et al., 2000). Width of the Reinforced Soil Zone, 11 m for Section 200, Approach Slab (3.72 m x 0.3 m) 75 mm Expanded Polystyrene Membrane & Collector Pipe Geogrid 1st layer Embedment Length is 8 m 0.3 m limit of 19 mm max. size crushed stone (3.81 m x 0.61 m) 2.055 m 2 9 R o w s f o r S e c ti o n 4 0 0 , 8 0 0 ( 5 .9 m h ig h ) 2 2 R o w s f o r S e c ti o n 2 0 0 (4 .5 m h ig h ) E m b e d m e n t 0 .4 5 m M in . F ro n t M S E W a ll Leveling Pad (0.15 m high) Block Unit (0.2 m high) CDOT Class 1 Backfill Connector Abutment Wall (0.76 m wide) Bridge Deck (0.13 m high) Girder (0.89 m high) Slope paving 1.35 m C a p U n it ( 0 .1 m h ig h ) 2 m Foundation 1.755 m Drainage Blanket with Pipe Drains 7.8 m Bedrock toward the top bottom with one to one slope increases linearly from 8 m at the The geogrid reinforcement length UX6 Geogrid UX6 Geogrid UX3 Geogrid Roadway (0.35 m high) Sleeper Foundation 0.4 m UX3 UX3 12.97 m for Sections 400 and 800 0.4 m high UX2

Case A5: Feather Falls Trail Bridge Abutments, California (Keller and Devin, 2003) A 12-m-long trail bridge was constructed in 1999 on the Feather Falls Trail in the Plumas National Forest in northern California. Because the project site was remote and without road access, the bridge materials had to be flown in with a heli- copter. Because of the deeply incised and narrow channel, the abutments were placed well above the channel high-water level. GRS abutments were selected for this project because they use small, lightweight materials and are easy to construct. Figure 2-7 shows the cross-section of the GRS abutments. The two abutments were 1.5 and 2.4 m high, and the wall facing comprised 0.15 m by 0.15 m treated timbers. Two polyester woven geotextiles of different strengths were used for the reinforcement. The top four layers of the reinforce- ments had an ultimate strength of 70 kN/m, while the remain- ing reinforcements had an ultimate strength of 52 kN/m, per ASTM D4595. The vertical reinforcement spacing was 0.15 m, and the average reinforcement length was 2.0 m. Most of the reinforcements were sandwiched and nailed between the fac- ing timbers, but the top four reinforcements were wrapped around the outside of the facing timbers and covered with timber boards to ensure maximum connection strength and to protect the geotextiles. Onsite rocky soil was used as the backfill and was com- pacted to 95 percent of its maximum dry density per AASHTO T-99. A geocomposite drain was placed behind each GRS 18 abutment, and each abutment had an embedment depth of 0.6 m to offer scour protection against a possible debris slide in the drainage. The entire construction of the bridge took about 2 weeks with a crew of two people. The GRS abutments have per- formed well since the bridge was put in service. Case A6: Alaska Bridge Abutments, Alaska (Keller and Devin, 2003) Two GRS abutments, constructed in 1992, support a 15.1-m-long precast, double-tee concrete bridge in the Tongass National Forest in southeast Alaska. Because trans- portation and construction costs are high in this area, the bridge and abutment designs had to be economical and easy to construct, without the need for specialized equipment. Because the bridge is located in the tidal-influence zone, there were concerns about corrosion loss, so GRS abutments were selected over hot-dipped galvanized welded-wire walls, which commonly had been adopted in the area. The GRS abutments were 3.7 m high and had three vertical faces: a front wall paralleled to the stream alignment and two wing walls oriented at 90 deg and 77 deg relative to the front-face wall. The distance between the front wall face to the toe of the sill was 0.9 m, and the distance between the centerline of the bearing of the bridge to the front wall face was 1.5 m. The combination of dead and live design loads caused by bridge superstructure was limited to 240 kPa. Figure 2-7. Cross-section of the Feather Falls Trail bridge abutments, California (Keller and Devin, 2003).

HDPE geogrids were used as the reinforcements. The base- to-height ratio for the front facing wall was 1:1 and was 0.7:1 for the two adjoining wing walls. The vertical spacing for the geogrid was 0.3 m near the base of the wall and 0.15 m near the top of the wall. Geogrids were wrapped around the timber facing, and 19 mm rebar drift pins were driven into pre-bored holes to hold the timber facing together. The full height of the wall was covered with 50-mm-thick treated timber board to protect the geogrids against UV degradation and floating debris. A free-draining granular material with maximum par- ticle size of 25 mm was used as the backfill. The GRS abut- ments have performed well since construction. Field Experiments of Bridge-Supporting Structures The test conditions and measured performance of six field experiments of GRS bridge abutments and piers are described below. The six field experiments are the Garden experi- mental embankment in France (Gotteland et al., 1997), the FHWA Turner-Fairbank GRS bridge pier in Virginia (Adams, 1997), the Havana Yard GRS bridge pier and abutment in Colorado (Ketchart and Wu, 1997), the Fiber Reinforced Plastic (FRP) geogrid-reinforced retaining wall in Japan (Miyata and Kawasaki, 1994), the Chemie Linz full-scale GRS embankment in Austria (Werner and Resl, 1986), and the Trento test wall in Italy (Benigni et al., 1996). 19 Case B1: Garden Experimental Embankment, France (Gotteland et al., 1997) A full-scale experiment was conducted in 1994 to investi- gate the failure behavior of GRS structures as bridge abut- ments, referred to as the “Garden” program (Geotextile: Application en Reinforcement, experimentation et Normali- sation). A 4.35-m-high embankment was constructed for the experiment, as shown in Figure 2-8. The test embankment was divided into two symmetrical parts corresponding to two different embankment profiles: the NW wall and the W wall. A fine sand used as backfill was compacted at its maximum standard Proctor density. The backfill had a dry unit weight of 16.6 kN/m3, friction angle of 30 deg, and cohesion of 2 kPa. Segmental concrete blocks were used as the facing. The NW wall was reinforced by a nonwoven geotextile with a tensile strength of 25 kN/m at 30 percent strain. The W wall was reinforced with a knitted woven geotextile with a tensile strength of 44 kN/m at 15 percent strain. The reinforcement spacing was 29 cm. The reinforced embankment was loaded in the same way as a bridge deck through a foundation slab. The 1.0-m-wide foundation was 1.5 m from the edge of facing. The embank- ment was loaded by a beam acted on by two thrust rams, each restrained by four tie-bars anchored into the embankment foundation. The test embankment was instrumented to monitor the performance of the embankment during loading. Two months Figure 2-8. The Garden experimental embankment, France (Gotteland et al., 1997). NW 2 sheets length = 3.6 m 2 sheets length = 3.6 m 1 m Geomembrane HDPE 4 .3 5 m NW Wall 1.5 m 38.41 m 4 m 4 m 5 m 3 m W 1 5 X 0 .2 9 m = 4 .3 5 m 1 m Intermediary Zone W Wall 1.5 m 4 m5 m

after the construction of the reinforced embankment, a load was applied to the foundation on the top of the structure until failure occurred. The load was applied over 2 days. The experiment was terminated when a permissible facing dis- placement was reached (0.20 m maximum horizontal dis- placement for the NW wall and 0.15 m for the W wall). The structure was examined layer by layer by careful excavation at the end of the experiment. A localized failure was noted to have occurred at the upper layers of the NW wall (with a nonwoven geotextile), whereas the W wall (with woven geotextile) experienced a deeper failure with a downstream tilting effect giving rise to a wide surface crack at the upstream end of the geotextile sheets. However, the main deformation occurred at the upper layers for both walls. The load-settlement curves for two walls showed a distinct break point that corresponds to two distinct slopes of the curves. The “critical loads” at the break point for the NW wall and the W wall were quite large, 140 kN/m and 123 kN/m respectively. The corresponding settlements were 36 mm and 33 mm, respectively. The lower “critical load” of the W wall can be attributed to its shorter “interme- diate” reinforcement (see Figure 2-8), even though the re- inforcement had higher strength than that of the NW wall. Case B2: FHWA Turner-Fairbank GRS Bridge Pier, Virginia (Adams, 1997) A full-scale bridge pier was constructed and load tested at the Turner-Fairbank Highway Research Center, FHWA, in McLean, Virginia. The pier was 5.4 m high and 3.6 m by 4.8 m at its base. The pier was supported on a reinforced soil foundation (RSF). The RSF comprised compacted road 20 base and three layers of biaxial geogrid reinforcement, spaced 0.3 m apart. The RSF was 1.2 m deep, over an area of 7.3 m by 7.5 m. Figure 2-9 shows the cross-section of the GRS bridge pier. The pier was constructed with modular concrete blocks as the facing and was reinforced with a polypropylene woven geotextile, Amoco 2044, at vertical spacing of 0.2 m. Because the geotextile was stronger in the cross-machine direction (38 kN/m at 5 percent strain) than in the machine direction (21 kN/m at 5 percent strain), the width and length direc- tions were alternated between layers. The backfill was clas- sified as a well-graded gravel. The maximum dry unit weight was 24 kN/m3, per AASHTO T-99, with the opti- mum moisture content being 5.0 percent. The average com- paction in the field was about 95 percent of the maximum dry density. The FHWA Turner-Fairbank pier was load-tested by applying vertical loads on top of the backfill in two loading cycles. The first loading cycle was performed when the pier height was 3.0 m. The 3.0-m-high pier was loaded to about 600 kPa. The settlement varied roughly linearly with the applied load. At 200 kPa, the settlement was about 13 mm, and at 600 kPa, the settlement was about 34 mm. The maxi- mum lateral displacements at 200 kPa and 600 kPa were about 6 mm and 20 mm, respectively. The second loading cycle was performed when the pier was at its full height. The second loading cycle was con- ducted in three parts: 1. The pier was incrementally loaded to 415 kPa and then held for 100 minutes; 2. The load was then ramped up to 900 kPa and held for 150 minutes and unloaded; and Figure 2-9. Cross-section of the FHWA Turner-Fairbank GRS bridge pier (Adams, 1997). Reinforced Soil Foundation Fabric Tail Between Blocks Compacted Road Base (Crushed Diabase) 0.2 m lifts Facing Block Leveling Pad (Woven Polypropylene Fabric) Reinforcement (Split Face Cinder Blocks) Dry Stacked Modular Blocks 5.4 m 1.2 m 3.8 m

3. The pier was reloaded to 415 kPa and then held for 100 minutes. At 415 kPa pressure, the pier settled about 25 mm; at 900 kPa, the settlement was about 70 mm. During the reload cycle, settlement was roughly reduced by a factor of two. At 200 kPa, the pier deformed laterally less than 3 mm. The maximum strain in the reinforcement was recorded near the middle of the pier and was about 2.3 percent. The reinforce- ment strain in the first loading cycle at 400 kPa was about 0.5 percent. Based on the measured results, the following conclusions were made: • At 200 kPa of loading, the GRS pier performed very sat- isfactorily. The maximum strain in the reinforcement was 0.25 percent. The maximum lateral displacement was 3 mm, yet no cracks occurred in the facing blocks. For the full-height load tests, the vertical settlement was about 15 mm in the initial load cycles and about 5 mm during the reload cycles. • Preloading reduced vertical settlement by about 50 per- cent and limited the vertical creep deformation. Pre- loading did not reduce lateral deformation. 21 • At 200 kPa, creep was not a concern in a closely spaced reinforced soil system with a well-compacted granular backfill. Case B3: Havana Yard GRS Bridge Pier and Abutment, Denver, Colorado (Ketchart and Wu, 1997) The Havana Yard GRS bridge supporting structures, con- sisting of two piers and one abutment, were constructed inside a 3.5-m-deep pit. The outer pier and the abutment were 7.6 m tall, and the center pier was 7.3 m tall. The center pier and the abutment were of a rectangular shape, and the outer pier was of an oval shape. The base of the outer pier, the cen- ter pier, and the abutment were, respectively, 2.4 m by 5.2 m (major and minor axes), 2.1 m by 4.8 m, and 3.6 m by 5.2 m. Segmental concrete blocks, each 0.2 m in height, were used as the facing element for all three structures. On the east face of the abutment, the facing assumed a 13 percent “negative batter” up to a height of 3.5 m. From 3.5 m to the top of the abutment were walking steps. Figure 2-10 shows the cross- section of the Havana Yard GRS bridge pier and abutment. The backfill was a “road base” material containing 13 per- cent of fines. The maximum dry unit weight was 21.2 kN/m3, Figure 2-10. Cross-section of the Havana Yard GRS bridge pier and abutment, Denver, Colorado, (Ketchart and Wu, 1997). Concrete Blocks Reinforcement spacing 0.3 m R e in fo rc e m e n t s p a c in g 0 .2 m R e in fo rc e m e n t s p a c in g 0 .2 m 7 .3 1 m 7 .6 2 m 7 .6 2 m Outer Pier Center Pier 5% 4%5% 4% 3% Girder R e in fo rc e m e n t s p a c in g 0 .2 m Abutment 13% 3 .5 3 m 0 .9 m 3% 2.44 m 0.74 m S te p s 5 .6 1 m

and the optimum moisture content was 6.7 percent, per AASHTO T-180. The field measured average dry unit weight was 19.1 kN/m3 (or 90 percent relative compaction) in the center pier and the abutment. The average placement moisture was 2.5 percent for the center pier and 1.6 percent for the abutment. The fill density of the outer pier was believed to be significantly lower than these measured val- ues because a lighter compaction plant was employed for the outer pier because of its size and shape. The reinforcement for all three structures was a woven polypropylene geotextile (Amoco 2044) with a wide width tensile resistance at 5 percent strain in the machine and cross- machine directions of 38 kN/m and 21 kN/m, respectively. The vertical spacing of the geotextile reinforcement was 0.2 m. The top four layers of the reinforcement in the abut- ment employed a wrapped-around procedure behind the facing block. A center geotextile “tail,” 1.2 m in length, was placed between each of these four layers to connect the back- fill to the facing blocks. On top of the outer pier and the abutment were 0.3-m-thick concrete pads to support steel bridge girders. The concrete pads were 0.9 m wide and 3.1 m long for the piers and 2.4 m wide and 3.7 m long for the abutment. The clearance distance of the concrete pad was 0.2 m from the back face of the fac- ing blocks. For loading tests, three steel bridge girders were placed over the top concrete pads of the outer pier and the abutment. Each girder was supported by steel bearing plates resting on the concrete pads. The steel bearing plates were located along the centerline of the top concrete pads. The span of the girders was 10.4 m. A total of 124 concrete blocks was placed on the girders. The total load was 2,340 kN, corre- sponding to an applied pressure of 232 kPa and 130 kPa on the outer pier and abutment, respectively. The findings and conclusions of this project as summa- rized by the authors are as follows: • The displacements of the pier and the abutment were comparable at applied load of 2,340 kN. The maximum vertical displacement was slightly higher in the outer pier than in the abutment. The maximum vertical dis- placements were 27.1 mm in the abutment and 36.6 mm in the outer pier, corresponding to 0.35 percent and 0.48 percent of the structure height. The maximum lateral displacement in the abutment was somewhat higher than that in the outer pier. The maximum lateral elon- gation of the perimeter was 4.3 mm in the abutment and 12.7 mm in the outer pier. • The ratio of the vertical movement to the structure height at 232 kPa of the outer pier (0.48 percent) was higher than that of the FHWA Turner-Fairbank pier (0.30 per- cent). This may be attributed to the much lower com- paction effort on the outer pier. The reinforcement strains in the fill direction of the outer pier and the FHWA Turner-Fairbank pier, however, were of similar 22 magnitude (0.2 percent to 0.4 percent). This suggests that the lateral movements of these piers are comparable. • Under a sustained load of 2,340 kN for 70 days, the creep displacements in both vertical and lateral direc- tions of the outer pier were about 4 times larger than those in the abutment because of lower compaction effort of the outer pier. The maximum vertical creep dis- placement was 61.6 mm in the outer pier and 18.3 mm in the abutment. The maximum lateral creep displace- ment was 59.5 mm in the outer pier and 14.3 mm in the abutment. • A significant part of the maximum vertical and lateral creep displacements of the pier and the abutment occurred in the first 15 days. At 15 days, the maxi- mum vertical and lateral creep displacements were about 70 percent to 75 percent of the creep displacements at 70 days in respective directions. • Creep deformation of the structures decreased with time. The vertical creep rates reduced nearly linearly (on log- log scale) with time. The creep rate of the outer pier (7.5 mm/day after 3 days and 0.1 mm/day after 70 days) was higher than that of the abutment (2.2 mm/day after 3 days and 0.03 mm/day after 70 days). • Hairline cracks of the facing blocks occurred in the outer pier and the abutment because of the lateral bulging and the down-drag force because of the friction between the backfill and the facing blocks. Installing flexible material (i.e., cushion) between vertically adjacent blocks may have alleviated this problem. • The maximum strains in the reinforcement were less than 1.0 percent. Compared with the rupture strain of the reinforcement of 18 percent, the safety margin against rupture of reinforcement appeared to be very high. • The calculated lateral displacements from the reinforce- ment strain distribution were in very good agreement with the measured lateral displacements. • With the less stringent construction condition (using a lightweight vibrating compaction plate), the outer pier showed about 1.5 times larger vertical displacement-to- height ratio than the Turner-Fairbank pier; whereas the lateral displacements were similar. Case B4: FRP Geogrid-Reinforced Retaining Wall, Japan (Miyata and Kawasaki, 1994) The FRP geogrid-reinforced test embankment consisted of three types of GRS retaining walls, referred to as Types A, B, and C (see Figure 2-11). The test embankment had a height of 5.0 m with a 0.3H:1V slope. Type A had a soft wall face with gabions only, Type B had a cement-treated wall face, and Type C was a gravity retaining wall made of cement-treated soil. An FRP geogrid, having a tensile strength of 49 kN/m at 2 percent strain, was used as re- inforcement.

To perform the load test, a 2.3-m-wide loading frame was first placed on the top of the embankment at 1.0 m from the front wall face. Loads were then applied by inserting loading steel plates into the loading frame in steps up to a total weight of 590 kN (q = 127 kPa). The lateral displacement of Type A was much larger than those of Types B and C. The maximum values were about 40 mm, 25 mm, and 20 mm for Types A, B, and C, respectively. In addition, the deformation mode of Type A differed from Types B and C. Type A, with a soft wall face, showed a swell-out mode with the maximum lateral movement occurring at about the mid-height of the wall. Types B and C, with a rigid wall surface, showed a forward fall-down mode with the maximum movement occurring at the top of the wall. Case B5: Chemie Linz Full-Scale GRS Embankment, Austria (Werner and Resl, 1986) A multi-layered geotextile reinforced embankment was built in 1981 and had been exposed to 3 years of extreme cli- matic fluctuations and environmental influences by the time it was loaded in 1984. The height of the embankment was 2.4 m. Figure 2-12 shows the geometry and loading scheme of the Chemie Linz full-scale GRS embankment. A silty gravelly sand was used as backfill. The design shear strength parame- ters of the backfill were φ = 21°, c = 20 kN/m2, and the bulk unit weight was 19.3 kN/m3. A polypropylene needle-punched nonwoven geotextile, with tear strength = 16 kN/m, grab strength = 1,200 N, and elongation at failure = 80 percent, as determined with DIN 53815 (German Standards: testing of textiles) was used as reinforcement. The vertical reinforce- ment spacing was 0.35 m. Wrapped facing was adopted for the structure. Seven steel slabs, each measuring 3 m by 1.3 m by 0.2 m, and two steel cylinders of 0.8 m in diameter were used to load the GRS embankment, which produced a total load of 23 510 kN or 130 kN/m2. Without rupture or critical defor- mation occurring, the load of 130 kN/m2 corresponded to 1.7 times the theoretical breaking load, as determined from Bishop’s lamella circular sliding surfaces method. The mea- sured maximum vertical settlement and lateral displacement of the embankment face were about 16 cm and 11 cm, respectively. Case B6: Trento Test Wall, Italy (Benigni et al., 1996) A 5-m-high test wall, referred to as the Trento test wall, was constructed in Northern Italy. A well-graded cohesion- less sandy gravelly soil, with shear strength parameters c´ = 100 kPa and φ´ = 40° determined from the CD triaxial tests, was used as backfill. It had a dry unit weight of 19.6 to 20.4 kN/m3 with in situ water content of 2.4 to 5.5 percent. Wrapped wall face was adopted for the wall. A geocompos- ite, with tensile strength of 27 kN/m at 16 percent strain per DIN EN ISO 10319 (German Standards: geotextiles wide- width tensile test), was used as reinforcement. The reinforced section of the wall was constructed in lifts separated by geo- composite layers and with the final spacing being 0.5 m. The reinforcement length was 2.0 m (40 percent of the wall height). Figure 2-13 shows the cross-section of the test wall. During construction, the wall face was supported by 1-m- high wooden forms, assembled with wide long boards nailed to brackets, which were wedged against a temporary scaffold. On completion of each lift, the underlying geosynthetic was wrapped around at the face and extended 2 m inside the back- fill. A new reinforced layer was then unrolled parallel to the wall face and positioned so that a 0.5-m-long tail rested on top of the one already wrapped around, while the remaining 2.5-m-long part draped over the wooden form. No windrows were used to anchor the reinforcement in the backfill. Figure 2-11. Cross-section of the FRP geogrid-reinforced retaining wall, Japan (Miyata and Kawasaki, 1994). 1 :0 .3 1 .5 m 1 .5 m 2 m 5 m Loading 1 :0 .3 Gabion FRP Geogrid (GB5) Separator Type A 2.3 m1 m Cement-treated soilCement-treated soil FRP Geogrid (GB5) Loading (GB10) 1 :0 .3 bearing capacity Reinforcements for Type B FRP Geogrid (GB5) Loading Type C

The loading test was performed by the weight of the stacked iron ingots evenly distributed over two 3 m by 3 m wide loading platforms placed on top of the wall. The maxi- mum surcharge loading, reached after 51 hours, was esti- mated at 84 kPa. The wall did not collapse under the applied load, although somewhat large movements were recorded. In addition, although most of the horizontal and vertical dis- placements were not recovered on unloading, it appeared that the wall had sustained almost no damage. Synthesis of Performance Characteristics The main performance characteristics of the 12 case his- tories reviewed in this study, including six in-service GRS bridge abutments and six full-scale field experiments, are 24 summarized in Table 2-1. The performance characteristics include wall height, backfill, reinforcement type, reinforce- ment spacing, facing type and connection, ratio of reinforce- ment length to wall height, maximum settlement of loading slab, maximum lateral movement of the wall face, maximum reinforcement strain, and failure pressure. Based on the measured performance, the following obser- vations are made in relation to performance, design, and con- struction of GRS bridge-supporting structures: • GRS bridge abutments with flexible facings are indeed a viable alternative to conventional bridge abutments. All six in-service GRS bridge abutments (Cases A1 through A6) exhibited satisfactory performance characteristics under service loads. The maximum settlements and max- imum lateral displacements for all the abutments were 1 .4 m 2 .4 m 1 m β= 8 5 ° 2.5 m β= 6 0 ° 1.3 m 2 .4 m (7  ~ 0 .3 5 m ) 9.8 m 3 m Slabs Steel Cylinders (a) Front View (b) Side View Figure 2-12. Geometry and loading scheme of the Chemie Linz full-scale GRS embankment, Austria (Werner and Resl, 1986).

under the tolerable movement criteria that were based on experience with real bridges—102 mm for settlement and 51 mm for lateral displacement (Grover, 1978; Bozozuk, 1978; Walkinshaw, 1978; Wahls,1990). • With a well-graded and well-compacted granular backfill and with closely spaced reinforcement (e.g., 0.2 m verti- cal spacing), the load carrying capacity of a GRS bridge supporting structure is very high (as high as 900 kPa in Case B2). The load-carrying capacity would be signifi- cantly smaller (e.g., 120 to 140 kPa in Case B1) if the backfill is of lower strength and the reinforcement is not of sufficient length (e.g., Case B1 where reinforcement extended only 0.3 m beyond the back edge of the sill). • With a well-graded and well-compacted granular backfill, the maximum settlement of the loading slab and the max- imum lateral movement of the wall face are very small under service loads (e.g., Cases A1, A4, and B2). With a lower quality backfill (as in Case B5, where the backfill was a silty gravelly sand with c = 20 kPa and φ = 21 deg and in Case A2, where the backfill was a fine sand with φ = 32 deg), the displacements would be significantly larger. • Fill placement density seems to play a major role in the performance of the GRS structures. For instance, Case B3 experienced 50 percent larger settlement than Case B2, even though the two GRS piers used the same reinforcement and the same reinforcement spacing. The difference in settlement resulted primarily from the difference in fill placement density and fill type. • Preloading can significantly reduce post-construction settlement of a GRS abutment (as in Cases A3 and B2) 25 by a factor of 2 to 6, depending on the initial placement density. In situations where the foundation soil is of dif- ferent thickness, preloading is an effective means to reduce differential settlement (as in Case A3). • With a well-graded and well-compacted granular back- fill, long-term creep under service loads can be negligi- bly small, as evidenced by Cases A4 and B2. • The maximum tensile strains in the reinforcement were in the range of 0.1 percent to 1.6 percent under service loads, with larger maximum strains being associated with lower strength backfill (e.g., 1.6 percent maximum strain in Case A2). • Reinforcement length and reinforcement type appeared to have only secondary effect on the performance char- acteristics. • The “sill clearance distance” (i.e., the distance between front edge of sill and back face of wall facing) employed in the cases vary fairly widely, from 0.2 m in Case B3 to 2.2 m in Case A2. A larger sill clearance will result in a longer bridge deck, thus higher costs, and may com- promise stability if the reinforcement is not sufficiently long (e.g., Case B1). THE NCHRP FULL-SCALE EXPERIMENTS Two full-scale experiments of segmental GRS bridge abutments, as shown in Figure 2-14, were conducted at the Turner-Fairbank Highway Research Center in McLean, Vir- ginia. The purposes of the full-scale experiments were to (1) examine the behavior of segmental GRS abutments subject to various load levels and (2) furnish a complete set of data (including material properties, material placement condi- tions, loading history, and measured and observed behavior) for verification of the analytical model employed in this study. Description of Test Sections The full-scale bridge abutments in the experiments con- sisted of two test sections. The two test sections were in a back- to-back configuration, as shown in Figure 2-15. The abutments were 4.65 m tall. Each test section had four components: (1) an abutment wall, (2) two wing walls, (3) a GRS mass, and (4) a sill on the top surface of the GRS mass near the edge of the wall facing. The geometry of the back-to-back test sections and the loading mechanism are shown in Figure 2-15. The back-to-back configuration had some advantages: (a) it elimi- nated the need to construct an approach fill or a retaining wall behind the abutment, thus reducing the amount of earthwork involved in the experiments; (b) it resulted in more consistent compaction of the fill across the two test sections—a key fea- ture to the success of the experiments; (c) it allowed the behav- ior of wing walls to be examined; (d) it avoided interferences Figure 2-13. Cross-section of the Trento test wall, Italy (Benigni et al., 1996). 0.5 m 0.6 m 5 m 0 .3 m 0 .5 m 2 m 1.5 m 0 2 6 4 8 3 m 10 0 .4 m (text continued on page 30)

26 TABLE 2-1 Case studies of GRS bridge-supporting structures with a flexible facing Case Height Backfill Reinf. Type Reinf. Spacing Facing Type and Facing Connection Reinf. Length to (Lower) Wall Height Ratio Maximum Settlement of Loading Slab Maximum Lateral Movement of Wall Face Maximum Reinf. Strain Failure Pressure Note Vienna Railroad Embankment (Case A1) 2.1 m c = 0 φ = 35° γ = 21 kN/m3 Needle-punched nonwoven geotextile with Tult = 23 kN/m @ε = 45% 0.3 m Wrapped face 0.8 1 mm under traffic load (60 kPa) Not reported Not reported Not loaded to failure New South Wales GRS Bridge Abutments (Case A2) 6.5 m and 9.5 m Compacted fine sand φ = 32° γdry(max) = 1.6 t/m3 (95% “standard relative density”) Tensar HDPE geogrid SR 110 with Tult = 110 kN/m @ε = 11.2% 0.4 m and 0.5 m Keystone blocks, with fiberglass dowels 1.2 to 1.6 80 mm @ service load 26 mm @ service load 1.6% @ service load Not loaded to failure Tiered (terraced) construction; sill clearance distance = 2.2 m Black Hawk Bridge Abutments (Case A3) 4.5 m and 7.5 m (lower wall) Silty clayey sand c = 34 kPa γdry = 15.8 kN/m3 (91% of T-99) w = 12% (2% dry of optimum) Amoco 2044, polypropylene woven geotextile with Tult = 70 kN/m @ε = 18% 0.3 m Natural rocks, with friction connection 0.7 to 1.2 Initial Loading: 4.9 to 28 mm @ 150 kPa; Reloading: 2.5 to 4.5 mm @ 150 kPa Initial Loading: 1.5 to 13 mm @ 150 kPa; Reloading: 0.6 to 4.5 mm @ 150 kPa 0.2% @ 80 kPa Not loaded to failure Preloading reduced differential settlements from 21.6 mm to less than 1.0 mm; sill clearance distance = 1.5 m Founders / Meadows Bridge Abutments (Case A4) 4.5 m and 5.9 m (lower wall) Gravelly sand γdry = 21 kN/m3 (95% of T-180) w = 5.6% (3.2% dry of optimum w/o gravels) Tensar HDPE geogrid UX6, UX3, & UX2, with Tult = 157, 64, & 39 kN/m, respectively 0.4 m Mesa concrete block, with plastic Mesa connectors 2.7 and 3.5 11 mm @ service load (150 kPa) after 18 months in service 13 mm @ service load (150 kPa) after 18 months in service 0.27 % after 33 months in service Not loaded to failure Small creep under service load; sill clearance distance = 1.35 m φ = 31°

27 TABLE 2-1 (Continued) Case Height Backfill Reinf. Type Reinf. Spacing Facing Type and Facing Connection Reinf. Length to (Lower) Wall Height Ratio Maximum Settlement of Loading Slab Maximum Lateral Movement of Wall Face Maximum Reinf. Strain Failure Pressure Note Feather Falls Trail Bridge Abutments (Case A5) 1.5 m and 2.4 m On-site rocky soil (95% of T-99) Polyester woven geotextiles with Tult = 52 and 70 kN/m 0.15 m Treated timber 1.3 and 0.8 Not reported Not reported Not reported Note loaded to failure Total cost = $320/m2 of wall face; no “bridge bump” Alaska Bridge Abutments (Case A6) 3.7 m Free-draining granular material with maximum particle size of 25 mm HDPE geogrid 0.3 m and 0.15 m Treated timber, with 19 mm rebar drift pins 1.0 and 0.7 Not reported Not reported Not reported Not loaded to failure Total cost = $452/m2 of wall face; no “bridge bump”; sill clearance distance = 0.9 m Garden Experimental Embankment (Case B1) 4.35 m Compacted fine sand c = 2 kPa φ = 30° γdry = 16.6 kN/m3 Nonwoven geotextile (Tult = 25 kN/m @ε = 30%) and woven geotextile (Tult = 44 kN/m @ε = 15%) 0.54 m (with “tails” for woven reinf.); 0.29 m below sill Concrete cells, with transverse synthetic bars 0.6 36 mm @ 140 kPa for nonwoven section; 33 mm @ 123 kPa for woven section Not reported 0.15% for nonwoven geotextile; 0.06% for woven geotextile Critical load = 140 kPa for nonwoven section (localized failure near the top); critical load = 123 kPa for woven section (deeper failure) Sill clearance distance = 1 m; other than the top two sheets, reinforcement only extended 0.3 m beyond the edge of sill

28 TABLE 2-1 (Continued) FHWA Turner- Fairbank GRS Bridge Pier (Case B2) 5.4 m Well-graded gravel γdry = 23 kN/m3 (95% of T-99) w = 3 to 7% (±2% of optimum) Amoco 2044, polypropylene woven geotextile with Tult = 70 kN/m @ε = 18% 0.2 m Cinder blocks, with friction connection 0.7 to 0.9 Initial Loading: 15 mm @ 200 kPa; 27 mm @ 415 kPa; 70 mm @ 900 kPa; Reloading: 8 mm @ 200 kPa; 13 mm @ 415 kPa Initial Loading: 3 mm @ 200 kPa; 9 mm @ 415 kPa; 35 mm @ 900 kPa; Reloading: 3 mm @ 200 kPa; 9 mm @ 415 kPa 2.3% @ 900 kPa Not loaded to failure Constructed with a well- compacted granular fill and small reinforcement spacing, the pier was loaded to 900 kPa without failure Case Height Backfill Reinf. Type Reinf. Spacing Facing Type and Facing Connection Reinf. Length to (Lower) Wall Height Ratio Maximum Settlement of Loading Slab Maximum Lateral Movement of Wall Face Maximum Reinf. Strain Failure Pressure Note Havana Yard GRS Bridge Pier and Abutment (Case B3) 7.6 m Road base material For abutment: dry = 19 kN/m3 (90% of T-180) w = 1.6% (5% of dry optimum) For pier: lower density than abutment Amoco 2044, polypropylene woven geotextile with Tult = 70 kN/m @ε = 18% 0.2 m Cinder blocks, with friction connection Abutment: 0.6 (typ.) Pier: 0.3 to 0.7 Abutment: 27 mm @ 130 kPa Pier: 37 mm @ 230 kPa Abutment: 14 mm @ 130 kPa Pier: 13 mm @ 230 kPa Abutment: 0.2% @ 130 kPa Pier: 0.4% @ 230 kPa Not loaded to failure Due to lower placement density, the pier experienced 50% larger settlement than the FHWA pier; sill clearance distance = 0.2 m γ

29 TABLE 2-1 (Continued) Case Height Backfill Reinf. Type Reinf. Spacing Facing Type and Facing Connection Reinf. Length to (Lower) Wall Height Ratio Maximum Settlement of Loading Slab Maximum Lateral Movement of Wall Face Maximum Reinf. Strain Failure Pressure Note FRP Geogrid- Reinforced Retaining Wall (Case B4) 5.0 m (H:V = 3:10) Not reported FRP (Fiberglass Reinforced Plastic) geogrid, with Tfail = 49 kN/m @ε = 2% FRP spacing = 1.5 to 2.0 m; tail spacing = 0.5 to 1.0 m; spacing below sill = 0.25 m Gabions (plastic bags filled with gravel) 0.9 Not reported 40 mm @ 130 kPa Not reported Not loaded to failure Very large reinforcement spacing (except directly beneath sill, with separate reinforcement for bearing capacity) Chemie Linz Full-Scale GRS Embankment (Case B5) 2.4 m Silty gravelly sand c = 20 kPa φ = 21° γ = 19.3 kN/m3 Polyfelt TS 400, polypropylene needle-punched nonwoven geotextile (Tult = 16 kN/m @ = 80%, weight = 350 g/m2) 0.35 m Wrapped face 1.0 160 mm @ 130 kPa 110 mm @ 130 kPa Not reported Not loaded to failure Structure loaded to 1.7 times the theoretical failure load; little creep Trento Test Wall (Case B6) 5.0 m Sandy gravelly soil c = 100 kPa φ = 40° 96 to 100% of T- 99 w = 2.2 to 5.3% dry of optimum Polyfelt PEC 50/25, a geocomposite with Tult = 27 kN/m @ε = 16% 0.5 m Wrapped face 0.35 50 mm @ 84 kPa 90 mm @ 130 kPa Not reported Not loaded to failure ε

30 from test sections to its right or left; and (e) it allowed the sill to be loaded without rotating in the longitudinal direction of the abutment because of uneven loading (if the test sections had been in a side-by-side configuration). The test abutments were constructed over a rigid floor. The rigid floor was a reinforced concrete mat measuring 9.1 m long, 7.3 m wide, and 0.9 m thick. The vertical spac- ing of the geosynthetic reinforcement for both test sections was 0.2 m in all layers. A concrete “cinder” block of dimen- sions 194 mm by 194 mm by 397 mm and with a split-face was used as a facing element. The front of each reinforce- ment sheet was placed between vertically adjacent facing blocks. No pins or any mechanical connectors were used between facing blocks. There was only frictional connection between the facing blocks and the reinforcement sheets. The length of all the reinforcements (“primary reinforcements”) was 3.15 m. In addition to the primary reinforcements, short “intermediate reinforcements” (1.3 m long) were placed at Figure 2-14. The NCHRP full-scale test abutment. 5.75 7.34 0.15 4.57 Sill Hydraulic jack 0.91 0.60 0.40 Wing wall Wing wall A b u tm e n t w a ll A b u tm e n t w a ll Section 2-2 Section 1-1 Load cell Hydraulic jack Sill Reaction plate Main loading beam Transverse loading beam Intermediate reinforcement sheet Facing blocks Reinforcement sheet Dywidag steel rod Strong concrete floor Anchor plate Units are in meters Mirafi test sectionAmoco test section Section 1-1 4.65 Section 2-2 Figure 2-15. Configuration of the NCHRP full-scale test abutments.

31 the mid-height of the top three facing blocks. These interme- diate reinforcements were placed immediately behind the facing block without connection to the facing. Placed on the top surface of each test section was a 0.3-m- thick concrete sill. The sill was 0.91 m wide and 4.57 m long, with its centerline aligned with the centerline of the abut- ment. The sill clear distance, measured from the back face of the abutment facing blocks to the front edge of the sill, was 0.15 m. The left and right edges of the sill were 0.40 m away from the back face of the wing walls. After the test abutment was constructed, a loading assem- bly was installed over the test section under the supervision of Michael Adams of the TFHRC. The loading assembly comprised a rigid floor (at the bottom of the assembly), hydraulic jacks (directly above the sill), steel rods (through the GRS mass), and reaction plates (at the top of the assem- bly). The rigid floor can accommodate up to 63 steel rods, each with a diameter of 44.5 mm (1.75 in.) with an allowable tensile load of 1,300 kN. Each steel rod was tied to an anchor base plate embedded in the rigid floor. Five steel rods were used for each test section. Vertical loads were applied on the sill through hydraulic jacks installed between the reaction plate and the sill. On applying hydraulic pressure to the jacks, the sill was pushed downward against the reaction plates and exerted vertical loads to the sill, hence the bridge abutment. A load cell was mounted between each hydraulic jack and the sill to monitor the applied loads. Construction Material and Placement Conditions Backfill The backfill was a non-plastic silty sand classified as SP-SM soil per USC system. The soil was considered representative of a “marginally acceptable” backfill for construction of GRS abutments. The soil has 8.5 percent of fine particles (passing the No. 200 sieve). The maximum dry unit weight of the soil was determined to be 18.3 kN/m3 with the optimum water content being 11.5 percent, per AASHTO T-99. The internal friction angle (φ) of the soil was 34.8 deg with a shear stress intercept (c) of 13.8 kPa. The shear strength parameters were determined by standard direct shear tests conducted on the part finer than the No. 10 sieve and prepared at 95 percent maximum dry unit weight, per AASHTO T-99. In the load tests, the target placement conditions were 100 percent compaction and ± 2 percent of the optimum moisture, per AASHTO T-99. Measurement taken after the load test showed that the compaction was 99.0 percent and the mois- ture was at 1.7 percent wet of optimum. With the information of the placement conditions of the fill, a series of large-size triaxial tests (with 150-mm-diameter, 300-mm-high speci- mens) in the “as-constructed” condition were conducted. The tests showed that the soil had φ = 37.3° and c = 20 kPa at the density and moisture mimicking the actual placement condi- tions. Large-size direct shear tests (with 300 mm by 300 mm specimens) conducted at the University of Massachusetts showed that the soil had φ = 36.5° and c = 0 kPa, tested in con- ditions mimicking the actual placement density and moisture. These soil property tests indicate that the fill is deemed acceptable by the current backfill selection criteria. Geotextile Reinforcement Everything was essentially the same for the two test sec- tions except for the geotextile reinforcement: one test section used Amoco 2044 (referred to as the Amoco test section) and the other used Mirafi 500x (referred to as the Mirafi test sec- tion). The wide-width tensile strength of Amoco 2044 and Mirafi 500x are Tult = 70 kN/m and Tult = 21 kN/m, respec- tively, in their cross-machine direction, per ASTM D 4595. Amoco 2044 was selected to represent a “lower bound” high- strength reinforcement, whereas Mirafi 500x represents a low- to medium-strength reinforcement. Both reinforcements are woven polypropylene geotextiles. Table 2-2 summarizes the main features of the two test sections with information on the test abutments’ configura- tion and the backfill and geotextile reinforcement properties. Construction of the NCHRP Test Abutments The construction procedure of the test abutments can be described by the following steps: 1. Level the surface of the rigid floor with a bedding sand; 2. Lay the first course of facing blocks to form a rectan- gular external dimension of 5.75 m by 7.34 m; 3. Place and compact backfill at the target density of 100 percent relative compaction using vibratory plate tampers; 4. Examine the field density by a nuclear density gauge; 5. Place two sheets of reinforcement, one in each test sec- tion, covering the top surface of the compacted backfill and the facing blocks; and 6. Lay the next course of facing blocks. Repeat Steps 3 to 5 until completion. Two different sizes of vibratory plate tampers were used in the construction. A lighter weight tamper (MBW AP- 2000, weighs 73 kg with a plate size of 48 cm by 53 cm) was used near the facing, whereas a heavier weight tamper (Mikasa MVH-304, weighs 315 kg with a plate size of 45 cm by 86 cm) was used in all other areas. Four to five passes were needed to achieve the targeted compaction. The construction of the two test sections began in mid- October 2002. On reaching a height of 1.2 m (i.e., with six courses of facing blocks), the construction had to be halted because of weather condition as described below.

32 The backfill of the test sections was placed at the pre- scribed density and moisture conditions, except for the last lift (wall elevation from 0.9 m to 1.2 m above base) wherein difficulties were encountered during fill com- paction. The lift was emplaced following a prolonged rainy day. The moisture content of the lift was in the range of 12.7 percent to 15.1 percent (i.e., 1.2 percent to 3.6 percent wet of optimum). In areas with high moisture contents (around 15 percent), the measured relative compaction was 95 percent per AASHTO T-99. Numerous attempts were made to increase the density by increasing the compaction passes of the vibrating tamper. Water, however, emerged from the top surface during the additional passes, and the measured density remained practically unchanged (relative compaction increased from 95.3 percent to 95.7 percent). Because of the weather, the construction had to be halted after some extended high-intensity precipitation at the con- struction site. Draining and drying of water from the back- fill did not appear possible absent an extended period of dry weather. In light of the difficulties with the placement density and moisture encountered in the 1.2-m-high abutments and the relatively “wet” winter experienced on the test site, it was judged necessary to remove the abutment and reconstruct two new test sections. The backfill on removal was found to be in a rather wet condition. There was also significant water accumulation near the base of the fill. Construction of the new GRS abutment test sections began in April 2003. The new test sections were constructed with the same backfill. There were concerns as to whether diffi- culties with placement density might be encountered as in the previous case. However, a decision was made to employ the same backfill so that the desired condition of using a “mar- ginally acceptable” backfill could be fulfilled. The top surface Amoco Test Section Mirafi Test Section Abutment height 4.65 m (15.25 ft) 4.65 m (15.25 ft) Sill 0.9 m × 4.5 m (3 ft × 15 ft) 0.9 m × 4.5 m (3 ft × 15 ft) Sill clear distance 0.15 m (6 in.) 0.15 m (6 in.) Reinforcement length 3.15 m (10 ft) 3.15 m (10 ft) Facing blocks (concrete) 194 mm x 194 mm × 397 mm (7.625 in. × 7.625 in. × 15.625 in.) 194 mm × 194 mm × 397 mm (7.625 in. × 7.625 in. × 15.625 in.) Vertical reinforcement spacing 0.2 m (8 in.) 0.2 m (8 in.) Reinforcements Amoco 2044: a woven polypropylene geotextile with T@ε = 1.0% = 12.3 kN/m (70 lb/in.) and Tult = 70 kN/m (400 lb/in.), per ASTM D4595, in the cross-machine direction. Mirafi 500x: a woven polypropylene geotextile with Tult = 21 kN/m (120 lb/in.), per ASTM D4595, in the cross-machine direction. Backfill For both test sections: a non-plastic silty sand (SP-SM, per USC System) Gradation: Percent passing 0.75-in. sieve = 100% Percent passing No.40 sieve = 59% Percent passing No.200 sieve = 8.5% Compaction Test, per AASHTO T-99: Maximum dry unit weight = 18.3 kN/m3 (116.5 lb/ft3) Optimum moisture content = 11.5% Standard Direct Shear Test (on the portion passing No. 10 or 2 mm sieve, at 95% maximum dry unit weight per AASHTO T-99; specimen size: 60 mm by 60 mm) Cohesion = 14 kPa (2 psi) Internal friction angle = 34.8° Large-size Direct Shear Test (at 99% maximum dry unit weight & 1.5% wet of optimum, per AASHTO T-99; specimen size: 300 mm by 300 mm) Cohesion = 0 kPa Internal friction angle = 36.5° Drained Triaxial Test (on the portion passing 9.5 mm or 3/8 in. sieve; at 99% maximum dry unit weight & 1.5% wet of optimum, per AASHTO T-99; specimen size: 150 mm diameter, 300 mm high) Cohesion = 20 kPa (3 psi) Internal friction angle = 37.3° TABLE 2-2 Main features of the NCHRP test abutments

33 of the fill was to be completely covered whenever there was any appreciable rainfall to better control the moisture. Backfill placement density and moisture were measured at the end of each construction lift to ensure that the specified val- ues were met. Four density tests were conducted for each lift. The average placement density and moisture were 99.0 per- cent compaction and 13.2 percent moisture for the Amoco test section and 98.4 percent compaction and 13.1 percent moisture for the Mirafi test section. Construction of the new test sections was completed near the end of May 2003. Instrumentation The instruments employed in the experiment included lin- ear voltage displacement transducers (LVDTs), displacement potentiometers, a laser displacement measurement device, strain gauges (for geotextile reinforcement), and contact pres- sure cells. Figure 2-16 shows the instrumentation layout for the experiments. For each test section, four displacement potentiometers and two LVDTs were used to measure settlements of the sill. Six displacement potentiometers were used to measure lateral movement of the abutment wall. In addition, six potentiome- ters and six LVDTs were used to measure lateral movements of the two wing walls. Three-dimensional movement of the abutment wall and one of the wing walls for each test section were traced by using a laser displacement measurement device. A total of 74 high-elongation strain gauges (Micro- Measurement Type EP-08-250BG-120) were mounted on five sheets of Amoco 2044 geotextile for the Amoco test section Figure 2-16. Layout of instrumentation. Instrumentation Layout of GRS Abutments 18 20 22 24 14 16 12 10 8 6 4 2 Potentiometer LVDT 2 4 6 8 10 12 14 16 18 20 22 24 LVDT Potentiometer Legend Layer F Layer H Layer I Layer JLayer E Layer D Layer C Layer B Layer A Strain gauge Concrete Foundation Contact pressure cell Layer G

34 and five sheets of Mirafi 500x geotextile for the Mirafi test sec- tion. The strain gauges were mounted along the centerline of the reinforcement sheets perpendicular to the abutment walls of the test sections. The locations of the strain gauges and rein- forcement sheets with strain gauges are shown in Figure 2-16. Three contact pressure cells (Geokon vibrating wire pressure transducer model 4810-25) were mounted on top of the rigid concrete foundation of the Mirafi test section. Loading The bridge sill was loaded along its centerline in equal increments of 50 kPa average vertical pressure. Each load increment was maintained for 30 minutes to allow the stress to be transferred to the entire soil mass. The first load test was carried out successfully on the Amoco test section on May 26, 2003. The loading was terminated at an average vertical pressure of 814 kPa, at which time the strokes of the loading rams had reached their maximum extension. The second load test was carried out successfully on the Mirafi test section on June 6, 2003. The loading was terminated at an average ver- tical pressure of 414 kPa, at which time the abutment expe- rienced “excessive” deformation. Measured Test Results and Discussions The results of loading tests on the Amoco and Mirafi test sections and the discussions of the test results are presented in this section. The test results reported include settlement of sill, lateral movement of abutment wall, tension crack in the soil mass, contact pressure on rigid foundation, and strains in geosynthetic reinforcements. Sill Settlement The sill settlements at six measured points including four corner points (with a legend “Pot”) and two mid-length points (with a legend “LVDT”) are shown in Figure 2-17 for the Amoco test section and in Figure 2-18 for the Mirafi test sec- tion. The measured data indicated that the front of the sill set- tled more than the back of the sill, while the left and right sides of the sill settled about the same. The average forward tilting of the Amoco and Mirafi test sections at 200 kPa pres- sure were about B/90 and B/50 (B = width of sill = 0.91 m), respectively. The average sill settlements of the six measurement points versus applied loads for both test sections are shown in Fig- ure 2-19. As expected, the settlement increased as the applied load increased. At a pressure of 200 kPa (the limiting bear- ing capacity of reinforced soil mass of an MSE abutment as recommended by the NHI manual, Elias et al., 2001), the average sill settlement in the Amoco test section was 40 mm, whereas the average sill settlement in the Mirafi test section was 72 mm. As the loading was terminated, 814 kPa for the Amoco test section and 414 kPa for the Mirafi test section, the average sill settlements were 163 mm and 175 mm, respectively. The Mirafi test section at 414 kPa had approached a bearing failure condition while the Amoco test Figure 2-17. Sill settlement versus applied pressure relationships of the Amoco test section. 0 20 40 60 80 100 120 140 160 180 200 0 100 200 300 400 500 600 700 800 900 1000 Applied Pressure (kPa) S il l S e tt le m e n t (m m ) Pot-PV-6 Pot-T-1 Pot-T-2 Pot-T-3 LVDT-CH-0 LVDT-CH-1

35 Figure 2-18. Sill settlement versus applied pressure relationships of the Mirafi test section. Figure 2-19. Average sill settlement versus applied pressure relationships of the Amoco and Mirafi test sections. 0 50 100 150 200 250 300 0 50 100 150 200 250 300 350 400 450 500 S il l S e tt le m e n t (m m ) Applied Pressure (kPa) Pot-PV-6 Pot-T-1 Pot-T-2 Pot-T-3 LVDT-CH-6 LVDT-CH-7 0 50 100 150 200 250 0 100 200 300 400 500 600 700 800 900 Applied Load (kPa) S il l S e tt le m e n t (m m ) Amoco Test Mirafi Test

36 section appeared to be sufficiently stable at 814 kPa. The Amoco and Mirafi test sections are essentially the same in all aspects except the reinforcement type. The difference in the sill settlement can be considered as a result of the difference in reinforcement stiffness and strength, Tult = 70 kN/m versus Tult = 21 kN/m. Lateral Wall Movement Figures 2-20 and 2-21 show the lateral movements of the abutment wall and wing-wall, respectively, of the Amoco test section. For the abutment wall, the maximum lateral movement occurred near the top of the wall (the top mea- surement point was not at the very top of the wall). The top one-third of the wall deformed at a much greater rate than the lower two-thirds of the wall. The maximum lateral movement was 24 mm and 82 mm at 200 kPa and 814 kPa, respectively. For the wing-wall, the lateral movements were much smaller than those of the abutment wall, with the maximum movement occurring at about H/6 (H = wall height) from the top of the wall for all the loads. The max- imum lateral movement was 18 mm and 33 mm at 200 kPa and 814 kPa, respectively. Figures 2-22 and 2-23 show the lateral movements of the abutment and wing-walls, respectively, of the Mirafi test section. For the abutment wall, the maximum lateral move- ment also occurred near the top of the wall under smaller loads. As the applied pressure exceeded about 300 kPa, the point of maximum wall movement shifted to H/6 from the top. Contrary to what was observed in the Amoco test section, the upper one-third of the Mirafi test section deformed at a slower rate than the lower two-thirds of the wall. At 200 kPa and 414 kPa, the maximum lateral movements were 36 mm and 115 mm, respectively. For the wing wall, the maximum lateral movement also occurred at about H/6 from the top of the wall. At 200 kPa and 413 kPa, the maximum lateral movements were 30 mm and 86 mm, respectively. The facing blocks in the top three courses were pushed outward as the sill tilted forward toward to wall face under higher applied loads. This suggests that (1) the sill clear distance of 0.15 m, a minimum value stipulated by the NHI manual, may be too small; and (2) it might be beneficial to increase the connection strength in the top three to four courses of the facing. The authors believe that it would be most effective to inter-connect the top three to four courses of the facing blocks after the construction is completed (i.e., after the deformation because of soil self-weight has occurred). Tension Cracks A tension crack on the wall crest was detected in both load tests when the average applied pressure on the sill was about 150 to 200 kPa. For both test sections, the tension crack was Figure 2-20. Lateral movement of abutment wall: Amoco test section. 0 0.5 1 1.5 2 2.5 3 3.5 4 4.5 5 0 10 20 30 40 50 60 70 80 90 Abutment-Wall Movement (mm) W a ll H e ig h t (m ) 53 kPa 98 kPa 160 kPa 207 kPa 260 kPa 312 kPa 371 kPa 424 kPa 475 kPa 581 kPa 685 kPa 733 kPa 814 kPa

37 Figure 2-21. Lateral movement of wing wall: Amoco test section. Figure 2-22. Lateral movement of abutment wall: Mirafi test section. 0 0.5 1 1.5 2 2.5 3 3.5 4 4.5 5 0 10 20 30 40 50 60 70 80 Wing-Wall Movement (mm) W a ll H e ig h t (m ) 53 kPa 98 kPa 160 kPa 207 kPa 260 kPa 312 kPa 371 kPa 424 kPa 475 kPa 581 kPa 685 kPa 733 kPa 814 kPa 0 0.5 1 1.5 2 2.5 3 3.5 4 4.5 5 0 20 40 60 80 100 120 140 Abutment-Wall Movement (mm) W a ll H e ig h t (m ) 53 kPa 101 kPa 164 kPa 214 kPa 263 kPa 317 kPa 375 kPa 414 kPa

38 about 11 m from the wall face, the location where the geosyn- thetic reinforcement ended. The cracks were parallel to the abutment wall face and extended through the entire width of the abutment. If an upper wall had been constructed over the test abutment, as in the case of typical bridge abutments, these tension cracks would not have been visible and perhaps less likely to occur. Hairline cracks of the facing blocks were also observed under higher loads in both test sections. Strains in Reinforcement A total of 74 strain gauges were mounted on five sheets of Amoco 2044 woven geotextile and five sheets of Mirafi 500x woven geotextile. The strain gauges were mounted by a “patch” technique. A strain gauge was first glued on the sur- face of a 25 mm by 76 mm patch. The patch was a low-strength heat-bonded nonwoven geotextile. To avoid inconsistent local stiffening of the patch because of the adhesive (given that the adhesive is much stiffer than the geotextile), the glue was applied only around the two ends of the strain gauge. The patch with a strain gauge already mounted was then glued on the reinforcement used in the experiments at a pre- scribed location, again with the glue applied only at the two ends. To protect the strain gauges from soil moisture and from possible mechanical damage during soil compaction, a microcrystalline wax was applied over the gauge and covered with a Neoprene rubber patch. This “patch” tech- nique for mounting strain gauges on nonwoven geotextiles has been used successfully in several projects including an FHWA pier (Adams, 1997), a Havana Yard pier and abut- ment (Ketchart and Wu, 1997), and a Black Hawk abutment (Wu et al., 2001). Because of the presence of the geotextile patch, calibra- tion of the strain gauges was needed. A wide-width tensile test was performed to correlate the recorded strain (local strain from strain gauges) with actual strain (average strain from the MTS machine) of the reinforcement. Figure 2-24 shows the calibration curve of Amoco 2044 and Mirafi 500x specimens. Because of the very long time lapse (about 8 months) between strain gauge installation and actual loading experi- ments, only 13 out of the 74 gauges worked properly. The most likely cause was that the gauges were damaged by the lengthy delay between mounting of strain gauges and actual loading experiments. Figures 2-25 and 2-26 show the mea- sured reinforcement strain versus applied pressure of the Amoco test section and Mirafi test section, respectively. Because of the limited number of operable strain gauges, dis- tributions of strain along any reinforcement sheet cannot be reliably deduced. The operable gauges, however, indicated that the maximum strains at 200 kPa were about 2.0 percent and 1.7 percent in the Amoco and Mirafi test sections, respectively. Figure 2-23. Lateral movement of wing wall: Mirafi test section. 0 0.5 1 1.5 2 2.5 3 3.5 4 4.5 5 0 20 40 60 80 100 120 140 Wing-Wall Movement (mm) W a ll H e ig h t (m ) 53 kPa 101 kPa 164 kPa 214 kPa 263 kPa 317 kPa 375 kPa 414 kPa

39 Figure 2-24. Calibration curve for (a) Mirafi 500x and (b) Amoco 2044. y = 1.17x 0 0.5 1 1.5 2 2.5 3 0 0.5 1 1.5 2 2.5 Measured Strain from Strain Gauge (%) A v e ra g e S tr a in f ro m M T S M a c h in e ( % ) (a) y = 1.17x 0 1 2 3 4 5 6 7 0 1 2 3 4 5 6 Measured Strain from Strain Gauge (%) A v e ra g e S tr a in f ro m t h e M T S M a c h in e ( % ) (b)

40 Figure 2-25. Reinforcement strains in the Amoco test section. Figure 2-26. Reinforcement strains in the Mirafi test section. 0 0.5 1 1.5 2 2.5 3 0 50 100 150 200 250 Applied Load (kPa) R e in fo rc e m e n t S tr a in ( % ) B-2 C-1 C-2 D-2 D-3 D-4 E-1 E-2 0 0.5 1 1.5 2 2.5 3 0 50 100 150 200 250 300 350 400 450 Applied Load (kPa) R e in fo rc e m e n t S tr a in ( % ) J-4 J-5 J-6 J-7 J-8

41 Figure 2-27. Measured contact pressure distribution on rigid foundation. Contact Pressures on the Rigid Foundation Figure 2-27 shows the measured contact pressures under different applied pressures at three selected points on the rigid floor of the Mirafi test section. The three points are located along the centerline of the abutment and are 0.25 m, 1.72 m, and 3.13 m from the wall face. The largest contact pressure occurred at 0.25 m from the wall face, which is roughly underneath the front of the sill. As expected, the contact pressure reduced with increasing distance from the wall face. At an average applied pressure of 200 kPa, the contact pressures at the three points were 77 kPa, 40 kPa, and 12 kPa. The NHI manual (Elias et al., 2001) recommends a method based on the 2:1 distribution with a cutoff at wall face to calculate vertical stress in a soil mass because of a concentrated vertical load applied on a footing for external and internal stability assessment. The contact pressure on the rigid foundation as calculated by the method is uniform at any given depth. The calculated contact pressures at the base were 23 kPa, 46 kPa, 69 kPa, and 92 kPa at applied pressures of 100 kPa, 200 kPa, 300 kPa, and 400 kPa, respectively. The corresponding measured average contact pressures are 24 kPa, 45 kPa, 59 kPa, and 77 kPa. A comparison of these con- tact pressures suggests that (1) the actual contact pressure is higher near the wall face and decreases nearly linearly with the distance from the wall face, and (2) the computation method in the NHI manual yields roughly the “average” measured contact pressure although the calculated pressures are somewhat higher than the measured average values at higher applied pressures. Summary of Measured Results of the NCHRP Test Abutments A summary of the measured performance and observed behavior of the two full-scale test abutments is presented in Table 2-3. FINDINGS FROM THE ANALYTICAL STUDY The analytical study was conducted by using a finite ele- ment code, DYNA3D (Hallquist and Whirley, 1989), and its PC version, LS-DYNA. The use of DYNA3D requires a workstation (such as CRAY, VAX/VMS, SUN, SILICON GRAPHICS, or IBM RS/6000), while LS-DYNA requires only a personal computer. The two computer codes give essentially the same results although LS-DYNA offers more user-friendly interfaces and greater flexibility in preparing the input file. The analytical model is briefly described in Appendix B. The capability of DYNA3D/LS-DYNA for analyzing the performance of segmental facing GRS bridge abutments was evaluated. The evaluation included comparing the analytical results with measured data of five full-scale experiments, 0 20 40 60 80 100 120 140 160 0 0.5 1 1.5 2 2.5 3 3.5 Distance from Abutment Wall Face (m) C o n ta c t P re s s u re ( k P a ) 100 kPa-Measured 200 kPa-Measured 300 kPa-Measured 400 kPa-Measured

42 including (1) the spread footing experiments by Briaud and Gibbens (1994), (2) the spread footing experiments on rein- forced sands by Adams and Collin (1997), (3) the FHWA Turner-Fairbank GRS bridge pier by Adams (1997), (4) the “Garden” experimental embankment in France (Gotteland et al., 1997), and (5) the two full-scale GRS bridge abutment loading experiments conducted as part of this study (referred to as the NCHRP GRS abutment experiment). Very good agreements between the analytical results and the measured data (including measured performance and failure loads) were obtained. The analyses of the five full-scale experiments are pre- sented in Appendix C, available as NCHRP Web-Only Doc- ument 81. The findings of a parametric study and findings of performance analysis, all obtained by using the analyti- cal model, are presented in this chapter. The findings of performance analysis were used as the basis for the allow- able sill pressures in the recommended design procedure. Parametric Study Base Case Geometry, Material Properties, and Loading After the finite element code, DYNA3D, was satisfacto- rily verified, a parametric study was conducted to investi- gate performance characteristics of GRS bridge abutments and the approach fill. The performance characteristics, as affected by soil placement condition, reinforcement stiff- ness/strength, reinforcement spacing (varying from 20 to 60 cm), reinforcement truncation, footing (sill) width, and the clearance between the front edge of the footing and the back face of the wall facing were investigated. When analyz- ing the results, the settlement of footing, rotation of the foot- ing, lateral deformation of abutment wall, maximum shear stress levels in the GRS soil mass, ultimate load carrying capacity of the abutment, and potential failure mechanisms were emphasized. TABLE 2-3 Summary of measured performance and observed behavior of the NCHRP test abutments Amoco Test Section Mirafi Test Section Reinforcement Amoco 2044,Tult = 70 kN/m Mirafi 500x, Tult = 21 kN/m Upon termination of loading: Average Applied Pressure 814 kPa 414 kPa Sill Settlement (front) 175 mm (6.9 in.) 189 mm (7.4 in.) Sill Settlement (back) 152 mm (6.0 in.) 160 mm (6.3 in.) Max. lateral movement in abutment wall 82 mm @ 4.5 m from base 115 mm @ 4.5 m from base Max. lateral movement in wing wall 33 mm @ 3.8 m from base 86 mm @ 3.8 m from base At 200 kPa (limiting bearing capacity, per NHI manual) Applied Pressure (average) 207 kPa 214 kPa Sill Settlement (front) 45 mm (1.8 in.) 81 mm (3.2 in.) Sill Settlement (back) 35 mm (1.4 in.) 64 mm (2.5 in.) Max. lateral movement in abutment wall 24 mm @ 4.5 m from base 36 mm @ 4.5 m from base Max. lateral movement in wing wall 18 mm @ 3.8 m from base 30 mm @ 3.8 m from base Observed Behavior: • The sills in both tests tilted toward the abutment wall face (i.e. the front of the sill settled more than the back; Left and right sides of the sill settled evenly. • The abutment wall leaned forward with the maximum movement occurring near the top of the wall. The top three courses of facing blocks were pushed outward at higher loads. • The wing wall also leaned forward with the maximum movement occurring at approximately 1/6H from the top of the wall. • In both tests, tension cracks occurred parallel to the wall face and were located at end of the reinforcement. Tension cracks initiated around 150–200 kPa average applied pressure. • Most strain gauges were damaged by moisture due to the long delay of actual loading experiments; maximum strain at 200 kPa was about 2.0% • The measured contact pressure on the rigid foundation was larger in front and decreased linearly toward the back. The computation procedure in the NHI Manual yielded about the average value of the contact pressures at a given applied load.

43 The analytical results obtained from the parametric study served as the basis for establishing preliminary design guide- lines of GRS abutments. The maximum tolerable settlement and horizontal movement of bridges, as suggested by Moul- ton et al. (1985) and by others (as summarized in NCHRP Report 343), may be used as the performance limits when establishing the design guidelines. The “Base Case” geometry used in the parametric analy- sis is shown schematically in Figures 2-28 and 2-29. The dimensions and parameters of the base case, listed below, were kept constant for all cases of the parametric study unless otherwise stated. Base Case Dimensions (see Figure 2-28): • Segmental wall height: 4.6 m • Total GRS abutment height: 7.1 m • Concrete block dimensions: 28 cm wide (toe to heel), 20 cm high, 50 cm long • Sill width: 1.5 m • Sill clearance: 15 cm • Loading pad width: 80 cm • Geosynthetic spacing: 20 cm • Geosynthetic length: 5 m Base Case Parameters: • Geosynthetic stiffness: 530 kN/m • Soil internal friction angle: 34 deg (See Figure 2-30) Base Case Loading: • Uniformly distributed vertical load was applied to the 80-cm-wide loading pad and increased monotonically until failure or until 1000 kPa was reached, whichever occurred first. Effects of Geosynthetic Spacing To investigate the effect of vertical spacing between geosynthetic layers on the GRS bridge abutment, three dif- ferent spacings were used: 20 cm (Base Case), 40 cm, and 60 cm. Four parameters were thought important in describing the performance of a GRS abutment when subjected to superstructure loads. These parameters, termed herein “per- formance parameters,” are the vertical displacement at the abutment seat (where the girder load is applied), the hori- zontal displacement at the abutment seat, the maximum dis- placement of the segmental facing, and sill distortion. Figure 2-31 shows the effects of increasing spacing on the selected performance parameters. Figure 2-31a shows that the vertical displacement at the abutment seat increases with spacing increase. The increase in displacement becomes more significant as the applied pressure increases. At 200 kPa of applied pressure (moderate pressure), there is a 24 percent increase in vertical displacement at 40 cm spacing as compared with the base case with 20 cm spacing. An increase Figure 2-28. Configuration of the base case for the parametric analysis.

44 Figure 2-29. Three-dimensional representation of the base case. Figure 2-30. Stress-strain-volume change characteristics of soils used in the parametric analysis.

45 of 50 percent in vertical displacement was observed at 60 cm spacing as compared with the base case. Similar trends with similar increases were noted for the horizontal displacement at abutment seat (Figure 2-31b) and for the maximum lateral displacement of the segmental facing (Figure 2-31c). The distortion of the sill, as shown in Figure 2-31d, ranged from +0.1 degree at 20 cm spacing to +0.41 degree at 60 cm spac- ing (positive distortion = forward tilt). At an applied pressure of 200 kPa, the vertical and hor- izontal displacements of the abutment seat for the base case were 4.7 cm and 2.1 cm, respectively (Figure 2-31). Judging from the criteria that the vertical movement should not exceed 100 mm and the horizontal movement should not exceed 50 mm (Wahls, 1990), the values of ver- tical and horizontal movements associated with the base case at 200 kPa pressure were deemed acceptable. This suggests that the displacements of the abutment are unlikely to cause any damage to the bridge superstructures. Judging from the same criteria, the vertical and horizontal displacements of the abutment seat for the base case at 400 kPa are unacceptable (barely acceptable): the vertical dis- placement is 10.3 cm, and the horizontal displacement is 4.6 cm (Figure 2-31). The maximum displacement criterion suggested by Wahls (1990) was based on a comprehensive study of bridge move- ments reported by Moulton et al. (1985). In the study, mea- sured movements were evaluated for 439 abutments of which most were perched abutments. The study included assess- ment of which movements were regarded as tolerable and which were intolerable. The tolerability of the movement was judged qualitatively by the agency responsible for each bridge in accordance with the following criterion: “Move- ment is not tolerable if damage requires costly maintenance and/or repairs and a more expensive construction to avoid this would have been preferable.” Effects of Backfill Soil Type Three backfill soils with internal friction angles of φ = 34°, 37°, and 40° and relative compactions (RC) of 95 percent, 100 percent, and 105 percent, respectively, are used in the analysis to investigate the effects of backfill soil type on the performance of the GRS abutment. The soil parameters used in the analysis were deduced from triaxial test results con- ducted on numerous backfill materials (Duncan et al., 1980). Figure 2-30 shows the stress-strain behavior and the volumet- ric strain-axial strain behavior of the three soils. The study by Duncan et al. (1980) presented estimates of stress-strain- strength parameters and volumetric strain-axial strain parame- ters for various soil types and degrees of compaction. These estimates were made using the compilations of data taken from 135 different soil parameters. Using these data, conservative parameter values have been interpreted for the soils under var- ious types and degrees of compaction. The values of stress- strain-strength parameters and volumetric strain-axial strain parameters of 16 materials averaged from the aforementioned 135 materials were presented in the study. These parameters Figure 2-31. Effects of geosynthetic spacing.

46 are called conservative in that they are typical of the lower val- ues of strength and modulus and the higher values of unit weight for each soil type. Figure 2-32 shows the effects of backfill soil type, as sig- nified by φ, on the performance parameters for geosynthetic spacings of 20 cm. More favorable response is attained when using soil types that have higher stiffness and strength and lower deformations. At 200 kPa of applied pressure, the ver- tical displacement at the abutment seat decreased 23 percent when φ increased from 34° (base case) to 37° as indicated in Figure 2-32a. The vertical displacement decreased 35 per- cent when φ was increased from 34° to 40°. The effect of increasing φ on the horizontal displacement of the abutment seat was similar in trend but with smaller magnitudes as shown in Figure 2-32b. As shown in Figure 2-32c, at 200 kPa of applied pressure, the maximum lateral displacement of the segmental facing decreased roughly linearly with increasing φ, with a total reduction of 45 percent at φ = 40° as compared with the base case. The distortion of sill changed from +0.1° at φ = 34° to +0.04° at φ = 40° as shown in Figure 2-32d. Figure 2-33 shows the effects of backfill soil type on the performance parameters for geosynthetic spacing s = 40 cm. As shown in Figure 2-33a, for φ = 34°, s = 40 cm, and at an applied pressure of 200 kPa, the vertical displacement of the abutment seat is 5.8 cm, which is 24 percent greater than that corresponding to the base case (φ = 34°, s = 20 cm, Figure 2-32a). For φ = 37° and s = 40 cm, the vertical displacement at the abutment seat is 9 percent smaller than the base case. For φ = 40° and s = 40 cm, the vertical displacement at the abutment seat is 28 percent smaller than that of the base case. The horizontal displacement at the abutment seat (Figure 2-33b) and the maximum lateral displacement of the seg- mental wall (Figure 2-33c) closely follow the trend of the vertical displacement at the abutment seat. The distortion of sill changed from +0.23° at φ = 34° to +0.035° at φ = 40° as shown in Figure 2-33d. The response of the base case (φ = 34°, s = 20 cm, Figure 2-32) is very similar to the case of φ = 37° and s = 40 cm, indicating that a better soil compaction may substitute for closer spacing (to a certain extent). Effects of Geosynthetic Stiffness The effects of geosynthetic stiffness (E * t) on the perfor- mance of the GRS abutment is shown in Figure 2-34 for geosynthetic spacings of 20 cm, and in Figure 2-35 for s = 40 cm. The stiffness of the base case was assumed to be 530 kN/m. A lower stiffness of 53 kN/m and a higher stiffness of 5300 kN/m were used to investigate the effects of geosyn- thetic stiffness on performance parameters. Figure 2-34a shows that the vertical displacement of the abutment seat of the base case is 4.7 cm for an applied pres- sure of 200 kPa. This displacement is reduced 43 percent when the geosynthetic stiffness is increased to 5300 kN/m. On the other hand, a drastic increase of 252 percent in dis- placement is noted when the geosynthetic stiffness is reduced to 53 kN/m. The same trend is noted for the horizontal dis- placement of the abutment seat (Figure 2-34b) and for the maximum lateral displacement of the segmental wall (Figure 2-34c). The distortion of the sill for the base case is +0.1° Figure 2-32. Effects of backfill internal friction angle for s = 20 cm.

47 Figure 2-33. Effects of backfill internal friction angle for s = 40 cm. Figure 2-34. Effects of geosynthetic stiffness for s = 20 cm.

48 Figure 2-35. Effects of geosynthetic stiffness for s = 40 cm. (forward tilt) as shown in Figure 2-34d. This distortion becomes −0.1° (backward tilt) when the stiffness is increased to 5300 kN/m. The distortion corresponding to a geosyn- thetic stiffness of 53 kN/m is +1.67°. Figure 2-35a indicates that at an applied pressure of 200 kPa, s = 40 cm, and E * t = 530 kN/m, there is a 24 percent increase in vertical displacement of the abutment seat as com- pared with the base case (Figure 2-34a). For E * t = 5300 kN/m and s = 40 cm, there is a 34 percent reduction in the magnitude of the vertical displacement at the abutment seat as compared with the base case (Figure 2-34a). Similar trends, but with greater changes, are noted in Figure 2-35b for the horizontal displacement of the abutment seat and in Figure 2-35c for the maximum lateral displacement of the segmental wall. The dis- tortion of the sill ranged from +1.8° at E * t = 53 kN/m to –0.05° at E * t = 5300 kN/m as shown in Figure 2-35d. Effects of Sill Clear Distance Sill clear distances of 0 cm, 15 cm (base case), and 30 cm were used to investigate the effects of clearance on the GRS abutment. The effects of sill clear distance on the perfor- mance of the GRS abutment is shown in Figure 2-36 for geosynthetic spacing s = 20 cm and in Figure 2-37 for s = 40 cm. Figure 2-36a shows that the vertical displacement of the abutment seat of the base case is 4.7 cm for an applied pres- sure of 200 kPa. This displacement is reduced 20 percent when the clearance is reduced to 0 cm. On the other hand, an increase of 11 percent in displacement is noted when the clearance is increased to 30 cm. The same trend and magni- tude is noted for the horizontal displacement of the abutment seat (Figure 2-36b). For the maximum lateral displacement of the segmental wall (Figure 2-36c), the trend was similar but with smaller magnitudes. The distortion of the sill for the base case is +0.1° (forward tilt) as shown in Figure 2-36d. This distortion becomes −0.1° (backward tilt) when the clearance is reduced to 0 cm. The distortion corresponding to a clearance of 30 cm is +0.22°. Figure 2-37a indicates that at an applied pressure of 200 kPa, s = 40 cm and a clear distance of 0 cm, there is a 4 per- cent increase in vertical displacement of the abutment seat as compared with the base case (Figure 2-36a). For a clear dis- tance of 30 cm and s = 40 cm, there is a 37 percent increase in the magnitude of the vertical displacement at the abutment seat as compared with the base case (Figure 2-36a). Similar trends with comparable magnitudes are noted in Figure 2-37b for the horizontal displacement of the abutment seat and in Figure 2-37c for the maximum lateral displacement of the segmental wall. The distortion of the sill ranged from –0.02° at a clear distance of 0 cm to +0.33° at a clear distance of 30 cm as shown in Figure 2-37d. Figures 2-36 and 2-37 show that the performance of the GRS abutment caused by decreasing sill clear distance is counter-intuitive. Decreasing clear distance indicates that the applied pressure is closer to the segmental facing, thus, greater displacements of the segmental facing, and therefore greater displacements at the abutment seat are expected. This discrepancy may be attributed to the fact that when the clear

49 Figure 2-36. Effects of sill clear distance for s = 20 cm. Figure 2-37. Effects of sill clear distance for s = 40 cm.

50 distance is small, there will be more contribution, in terms of stiffness, from the segmental facing. Nevertheless, this counter-intuitive response can be ascertained via large-scale testing of a GRS bridge abutment with small and large sill clearances. Effects of Sill Width To investigate the effect of sill width on the GRS bridge abutment, two sill widths were used: 150 cm (Base Case), and 100 cm. The effects of sill width on the performance of the GRS abutment is shown in Figure 2-38 for s = 20 cm and in Figure 2-39 for s = 40 cm. At 300 kN/m of applied load (corresponding to 200 kPa of applied pressure for the 150-cm-wide sill, and 300 kPa for the 100-cm-wide sill), the vertical displacement at the abut- ment seat increased 21 percent when the width decreased from 150 cm (base case) to 100 cm, as indicated in Figure 2-38a. The effect of decreasing sill width on the horizontal displacement of the abutment seat was similar in trend and magnitude as shown in Figure 2-38b. As shown in Figure 2-38c, at 300 kN/m of applied load, the maximum lateral dis- placement of the segmental facing increased roughly 11 per- cent when sill width decreased to 100 cm. The distortion of sill changed from +0.1° at sill width of 150 cm to +0.18° at sill width of 100 cm as shown in Figure 2-38d. As shown in Figure 2-39a, for a sill width of 150 cm, s = 40 cm, and an applied load of 300 kN/m, the vertical dis- placement of the abutment seat is 5.8 cm, which is 24 percent greater than that corresponding to the base case (sill width = 150 cm, s = 20 cm, Figure 2-38a). For a sill width of 100 cm and s = 40 cm, the vertical displacement at the abutment seat is 67 percent greater than the base case. The horizontal displacement at the abutment seat (Figure 2-39b) and the maximum lateral displacement of the segmental wall (Figure 2-39c) follow the same trend of the vertical displacement at the abutment seat. The distortion of sill changed from +0.4° at sill width of 150 cm to +0.28° at sill width of 100 cm, as shown in Figure 2-39d. Figures 2-38 and 2-39 show that the performance parameters increased at a higher rate under higher applied loads. Effects of Reinforcement Truncation To study the effects of truncated reinforcement on the per- formance of the GRS abutment, the GRS abutment was mod- ified so that the reinforcement length is truncated at the base. The truncated base was assumed to be H/4 and increases upward at 45° angle (H is the height of the segmental wall). Figure 2-40 shows the effect of truncated reinforcement on the performance parameters for s = 40 cm. The figure com- pares the truncated and non-truncated reinforcement cases and indicates that the effect of truncated reinforcement is insignificant in terms of displacements. Figure 2-40d shows that there is a small decrease in sill distortion in the case of truncated reinforcement. Figure 2-38. Effects of sill width for s = 20 cm.

51 Figure 2-39. Effects of sill width for s = 40 cm. Figure 2-40. Effects of reinforcement truncation for s = 40 cm. Predicting Failure Loads In the parametric study, none of the GRS bridge abutments failed “catastrophically” as in the Garden experiment and the Garden test analysis described in Appendix C (NCHRP Web- Only Document 81). All abutments withstood the 1000 kPa load but suffered very significant displacements and dis- tresses without sudden failure. It is suitable to think about shear strain in the soil mass as a measure of distress in a GRS abutment. Thus, a simple failure criterion based on the max- imum shear strain is proposed herein in order to estimate the allowable bearing pressure of a spread footing.

52 From the parametric analysis, it was noted in most cases that there exists a triangular failure zone initiating near the soil-facing interface and propagating into the backfill as the load is increased. This failure zone sustained greater shear strains than the rest of the backfill. The proposed failure cri- terion suggests that failure occurs when (1) the triangular shear zone propagates all the way to the back edge of the spread footing as shown in Figure 2-41, and (2) the triangular shear zone sustains shear strains that exceed a critical value, γ(critical), defined as: γ(critical) = 2/3 [(1)failure – (3)failure] where (1)failure is the axial strain at failure, (3)failure is the radial strain at failure, and 2/3 is a reduction factor (equivalent to a safety factor of 1.5). Both (1)failure and (3)failure can be obtained from triaxial test results. For the soils used in the parametric analysis with φ = 34°, 37°, and 40°, the critical shear strain, γ(critical), is determined as 3.2 percent with the help of the triax- ial test results presented in Figure 2-30. This critical shear strain value was then used to estimate the allowable bearing pressure under different conditions as shown in Table 2-4. Load-Carrying Capacity Analysis A series of load-carrying capacity analyses, as an exten- sion of the Parametric Study presented in the previous sec- tion, was conducted to examine the effect of sill type, sill width, soil stiffness/strength, reinforcement spacing, and foundation stiffness on the allowable load-carrying capacity of GRS abutment sills. Seventy-two analyses were per- formed using the LS-DYNA code. The variables in the analyses included the following: • Sill type: integrated sill and isolated sill; • Sill width: 0.8 m, 1.5 m, and 2.5 m; • Reinforcement spacing: 0.2 m and 0.4 m; • Soil friction angle: 34°, 37°, and 40°; and • Foundation: 6-m-thick medium sand foundation and rigid foundation. Of the 72 analyses, one-half were for a GRS abutment situated over a medium sand foundation (with its stiffness representing a lower-bound “competent” foundation), and the other half were for a GRS abutment situated over a rigid foundation. Geometry and Material Properties Figures 2-42 through 2-47 show the configuration of the abutments investigated in this study. Sill widths of 0.8 m, 1.5 m, and 2.5 m, and two sill types (integrated and isolated sills) were investigated. The abutments share the following common features: • Lower wall height = 4.67 m (15.3 ft); upper wall height = 2.44 m (8 ft); Figure 2-41. The critical shear strain distribution failure criterion.

53 • Sill clear distance = 0.15 m (6 in.); • Same backfill in the upper and lower walls; • A conservative secant modulus for geosynthetic rein- forcements = 530 kn/m; and • Reinforcement length in lower wall = 5.0 m (0.7 * the total wall height); reinforcement length in upper wall = 7.5 m. The case of an abutment over a rigid foundation with an integrated sill (sill width = 1.5 m) has the same config- uration as the base case in the Parametric Study (see Fig- ure 2-28). The reinforced fill, retained earth, and the medium sand foundation were simulated by an extended two-invariant geologic cap material model. The material parameters of the geologic cap model for the three select backfills with fric- tion angles of 34°, 37°, and 40° and the retained earth behind the reinforced soil region are summarized in Table 2-5. The medium sand foundation was assumed to have the same properties as the reinforced fill with a friction angle of 37°. Also, the same stiffness values as those used in the Parametric Study for soils with φ = 34°, 37°, and 40° were employed. The sill, modular facing blocks, approach slab, and geosynthetic reinforcement were simulated by an elastic material model. The elastic material parameters are summa- rized in Table 2-6. The geosynthetic reinforcement in the performance analysis was assumed to have a constant stiff- ness (E * t) of 530 kN/m. The material parameters listed in Table 2-6 are of the same values as those used in the base case of the Parametric Study. Performance Characteristics The results of the 72 finite element analyses were summa- rized in 32 figures. Four performance characteristics were examined in the figures: settlement of sill, maximum lateral displacement of wall face, lateral movement of sill, and rota- tion of sill. Each of the 32 figures shows the relationships between the applied pressure on the sill and one of the four performance characteristics for three different sill widths (0.8 m, 1.5 m, and 2.5 m) and three soil friction angles (34°, 37°, and 40°). The conditions associated with each figure are as follows: • The relationship between settlement of sill and applied sill pressure: Figure 2-48 (s = 0.2 m, integrated sill), Fig- ure 2-49 (s = 0.4 m, integrated sill), Figure 2-50 (s = 0.2 m, isolated sill), Figure 2-51 (s = 0.4 m, isolated sill), all on the medium sand foundation. • The relationship between lateral displacement of wall face and applied sill pressure: Figure 2-52 (s = 0.2 m, integrated sill), Figure 2-53 (s = 0.4 m, integrated sill), Figure 2-54 (s = 0.2 m, isolated sill), Figure 2-55 (s = 0.4 m, isolated sill), all on the medium sand foundation. • The relationship between lateral movement of sill set- tlement and applied sill pressure: Figure 2-56 (s = 0.2 m, integrated sill), Figure 2-57 (s = 0.4 m, integrated sill), Figure 2-58 (s = 0.2 m, isolated sill), Figure 2-59 (s = 0.4 m, isolated sill), all on the medium sand foun- dation. • The relationship between rotation of sill and applied sill pressure: Figure 2-60 (s = 0.2 m, integrated sill), Figure 2-61 (s = 0.4 m, integrated sill), Figure 2-62 (s = 0.2 m, isolated sill), Figure 2-63 (s = 0.4 m, isolated sill), all on the medium sand foundation. • Figures 2-64 to 2-79 (16 figures) correspond to the same conditions as Figures 2-48 to 2-63 (also 16 fig- ures), except that the abutments are situated over a rigid foundation. The sill settlements and sill lateral movements presented in the figures have excluded the deformations caused by self- weight of the soil because those deformations in actual con- struction can and will be compensated for or adjusted to zero before any sill pressure is applied. On the other hand, the lat- eral maximum displacement of the wall face cannot be com- pensated for or adjusted to from the onset of construction, thus the accumulated values are reported. General observations of the performance characteristics follow: • For reinforcement spacing of 0.2 m, none of the abut- ments suffered from any stability problems up to an applied pressure of 1,000 kPa. φ = 34° Reinforcement Spacing = 20 cm 225 kPa 280 kPa 360 kPa Reinforcement Spacing = 40 cm 120 kPa 200 kPa 280 kPa φ = 37° φ = 40° TABLE 2-4 Allowable bearing pressures based on the critical shear strain distribution criterion (text continues on page 60)

54 Figure 2-42. Configuration of a GRS abutment with integrated sill, sill width = 0.8 m.

55 Figure 2-43. Configuration of a GRS abutment with integrated sill, sill width = 1.5 m.

56 Figure 2-44. Configuration of a GRS abutment with integrated sill, sill width = 2.5 m.

57 Figure 2-45. Configuration of a GRS abutment with isolated sill, sill width = 0.8 m.

58 Figure 2-46. Configuration of a GRS abutment with isolated sill, sill width = 1.5 m.

59 Figure 2-47. Configuration of a GRS abutment with isolated sill, sill width = 2.5 m.

60 φ = 34° soil φ = 37° soil φ = 40° soil Retained Earth(φ = 30°) Initial bulk modulus, K (MPa) 16.45 24.67 32.89 16.45 Initial shear modulus, G (MPa) 7.59 11.39 15.18 7.59 Failure envelope parameter, α (kPa) 0 0 0 Failure envelope linear coefficient, θ 0.264 0.289 0.315 0.231 Cap surface axis ratio, R 4 4 4 4 Hardening law exponent, D (kPa)−1 7.25 × 10−6 7.25 × 10−6 7.25 × 10−6 7.25 × 10−6 Hardening law coefficient, W 2.5 1.5 1.0 2.5 Hardening law parameter, X0 (kPa) 200 200 200 0 0 Elastic Material Parameters Sill, Facing blocks, and Approach slab Geosynthetic reinforcement Young’s modulus, E (kPa) 13.8 × 106 4.14 × 104 Poisson’s ratio, ν 0.21 0.3 • For reinforcement spacing of 0.4 m, most of the abut- ments encountered facing failure (i.e., the top two to three courses of facing blocks “fell off” the wall face) when the applied pressure exceeded 500 kPa to 970 kPa (depending on the geometric condition and material properties of the abutment). Only those abutments with sill width = 0.8 m and soil friction angle = 37° and 40° did not encounter facing failure up to an applied pressure of 1,000 kPa. In any case, there was no cata- strophic failure in any abutment up to 1,000 kPa applied pressure. • For reinforcement spacing of 0.2 m, the rate of de- formation was relatively small at an applied pres- sure between 0 to 100 kPa. The rate of deformation increased slightly between 100 to 200 kPa and then remained roughly constant between 200 and 1,000 kPa. • For reinforcement spacing of 0.4 m, the rate of defor- mation was also relatively small at an applied pressure between 0 to 100 kPa. For applied pressure between 100 and 400 kPa, the rate of deformation increased somewhat with increasing pressure. Once exceeding about 400 kPa, the rate of deformation was nearly constant until it approached a failure condition (facing failure). • The settlement of sill was somewhat less for the abut- ments with an integrated sill than with an isolated sill; whereas the maximum lateral displacement of wall face did not differ much for the two types of sill. • The differences in the magnitude of the performance characteristics for φ between 34° and 37° were gener- ally greater than those between 37° and 40°. This suggests that increasing the soil friction angle (by selecting a better fill type and/or with better compaction efforts) to improve the performance characteristics is more efficient for soils with a lower friction angle than for soils with a higher friction angle. • The effect of reinforcement spacing on sill settlement and maximum lateral displacement of wall face was significant, especially at applied pressure greater than 200 kPa. TABLE 2-5 Geologic cap model material parameters TABLE 2-6 Elastic model material parameters (text continues on page 93)

61 Figure 2-48. Relationship between applied pressure and sill settlement: integrated sill, s= 0.2 m, and medium sand foundation.

62 Figure 2-49. Relationship between applied pressure and sill settlement: integrated sill, s = 0.4 m, and medium sand foundation.

63 Figure 2-50. Relationship between applied pressure and sill settlement: isolated sill, s = 0.2 m, and medium sand foundation.

64 Figure 2-51. Relationship between applied pressure and sill settlement: isolated sill, s = 0.4 m, and medium sand foundation.

65 Figure 2-52. Relationship between applied pressure and maximum lateral wall displacement: integrated sill, s = 0.2 m, and medium sand foundation.

66 0 2 4 6 8 10 12 0 100 200 300 400 500 600 700 800 900 1000 Applied Pressure (kPa) M a x i m u m L a t e r a l W a l l M o v e m e n t / L o w e r W a l l H e i g h t ( % ) denotes facing failure (facing failure @ 817 kPa) (facing failure @ 667 kPa) (facing failure @ 567 kPa) (facing failure @ 917 kPa) (facing failure @ 767 kPa) (facing failure @ 650 kPa) φ = 37° (Sill Width = 1.5 m) φ = 40° (Sill Width = 1.5 m) φ = 34° (Sill Width = 2.5 m) φ = 37° (Sill Width = 2.5 m) φ = 40° (Sill Width = 2.5 m) (facing failure @ 969 kPa) φ = 37° (Sill Width = 0.8 m) φ = 40° (Sill Width = 0.8 m) φ = 34° (Sill Width = 1.5 m) φ = 34° (Sill Width = 0.8 m) Figure 2-53. Relationship between applied pressure and maximum lateral wall displacement: integrated sill, s = 0.4 m, and medium sand foundation.

67 0 2 4 6 8 10 12 0 100 200 300 400 500 600 700 800 900 1000 Applied Pressure (kPa) M a x i m u m L a t e r a l W a l l M o v e m e n t / L o w e r W a l l H e i g h t ( % ) φ = 34° (Sill Width = 0.8 m) φ = 37° (Sill Width = 0.8 m) φ = 40° (Sill Width = 0.8 m) φ = 34° (Sill Width = 1.5 m) φ = 37° (Sill Width = 1.5 m) φ = 40° (Sill Width = 1.5 m) φ = 34° (Sill Width = 2.5 m) φ = 37° (Sill Width = 2.5 m) φ = 40° (Sill Width = 2.5 m) Figure 2-54. Relationship between applied pressure and maximum lateral wall displacement: isolated sill, s = 0.2 m, and medium sand foundation.

68 0 2 4 6 8 10 12 0 100 200 300 400 500 600 700 800 900 1000 Applied Pressure (kPa) M a x i m u m L a t e r a l W a l l M o v e m e n t / L o w e r W a l l H e i g h t ( % ) φ = 34° (Sill Width = 0.8 m) (facing failure @ 937 kPa) φ = 34° (Sill Width = 1.5 m) φ = 40° (Sill Width = 0.8 m) φ = 37° (Sill Width = 0.8 m) denotes facing failure (facing failure @ 717 kPa) (facing failure @ 617 kPa) (facing failure @ 533 kPa) (facing failure @ 867 kPa) (facing failure @ 750 kPa) (facing failure @ 633 kPa) φ = 40° (Sill Width = 2.5 m) φ = 37° (Sill Width = 2.5 m) φ = 34° (Sill Width = 2.5 m) φ = 40° (Sill Width = 1.5 m) φ = 37° (Sill Width = 1.5 m) Figure 2-55. Relationship between applied pressure and maximum lateral wall displacement: isolated sill, s = 0.4 m, and medium sand foundation.

69 -1 0 1 2 3 4 5 6 7 8 9 10 0 100 200 300 400 500 600 700 800 900 1000 Applied Pressure (kPa) S i l l L a t e r a l M o v e m e n t / L o w e r W a l l H e i g h t ( % ) φ = 34° (Sill Width = 0.8 m) φ = 40° (Sill Width = 2.5 m) φ = 37° (Sill Width = 2.5 m) φ = 34° (Sill Width = 2.5 m) φ = 40° (Sill Width = 1.5 m) φ = 37° (Sill Width = 1.5 m) φ = 34° (Sill Width = 1.5 m) φ = 40° (Sill Width = 0.8 m) φ = 37° (Sill Width = 0.8 m) Figure 2-56. Relationship between applied pressure and sill lateral movement: integrated sill, s = 0.2 m, and medium sand foundation.

70 -1 0 1 2 3 4 5 6 7 8 9 10 0 100 200 300 400 500 600 700 800 900 1000 Applied Pressure (kPa) S i l l L a t e r a l M o v e m e n t / L o w e r W a l l H e i g h t ( % ) denotes facing failure (facing failure @ 817 kPa) (facing failure @ 667 kPa) (facing failure @ 567 kPa) (facing failure @ 917 kPa) (facing failure @ 767 kPa) (facing failure @ 650 kPa) φ = 37° (Sill Width = 1.5 m) φ = 40° (Sill Width = 1.5 m) φ = 34° (Sill Width = 2.5 m) φ = 37° (Sill Width = 2.5 m) φ = 40° (Sill Width = 2.5 m) (facing failure @ 969 kPa) φ = 37° (Sill Width = 0.8 m) φ = 40° (Sill Width = 0.8 m) φ = 34° (Sill Width = 1.5 m) φ = 34° (Sill Width = 0.8 m) Figure 2-57. Relationship between applied pressure and sill lateral movement: integrated sill, s = 0.4 m, and medium sand foundation.

71 -1 0 1 2 3 4 5 6 7 8 9 10 0 100 200 300 400 500 600 700 800 900 1000 Applied Pressure (kPa) S i l l L a t e r a l M o v e m e n t / L o w e r W a l l H e i g h t ( % ) φ = 34° (Sill Width = 0.8 m) φ = 37° (Sill Width = 0.8 m) φ = 40° (Sill Width = 0.8 m) φ = 34° (Sill Width = 1.5 m) φ = 37° (Sill Width = 1.5 m) φ = 40° (Sill Width = 1.5 m) φ = 34° (Sill Width = 2.5 m) φ = 37° (Sill Width = 2.5 m) φ = 40° (Sill Width = 2.5 m) Figure 2-58. Relationship between applied pressure and sill lateral movement: isolated sill, s = 0.2 m, and medium sand foundation.

72 -1 0 1 2 3 4 5 6 7 8 9 10 0 100 200 300 400 500 600 700 800 900 1000 Applied Pressure (kPa) S i l l L a t e r a l M o v e m e n t / L o w e r W a l l H e i g h t ( % ) φ = 34° (Sill Width = 0.8 m) (facing failure @ 937 kPa) φ = 34° (Sill Width = 1.5 m) φ = 40° (Sill Width = 0.8 m) φ = 37° (Sill Width = 0.8 m) denotes facing failure (facing failure @ 717 kPa) (facing failure @ 617 kPa) (facing failure @ 533 kPa) (facing failure @ 867 kPa) (facing failure @ 750 kPa) (facing failure @ 633 kPa) φ = 40° (Sill Width = 2.5 m) φ = 37° (Sill Width = 2.5 m) φ = 34° (Sill Width = 2.5 m) φ = 40° (Sill Width = 1.5 m) φ = 37° (Sill Width = 1.5 m) Figure 2-59. Relationship between applied pressure and sill lateral movement: isolated sill, s = 0.4 m, and medium sand foundation.

73 Figure 2-60. Relationship between applied pressure and rotation of sill: integrated sill, s = 0.2 m, and medium sand foundation. -1 0 1 2 3 4 5 0 100 200 300 400 500 600 700 800 900 1000 Applied Pressure (kPa) R o t a t i o n o f S i l l ( ° ) , P o s i t i v e f o r C o u n t e r - C l o c k w i s e R o t a t i o n φ = 34° (Sill Width = 0.8 m) φ = 40° (Sill Width = 2.5 m) φ = 37° (Sill Width = 2.5 m) φ = 34° (Sill Width = 2.5 m) φ = 40° (Sill Width = 1.5 m) φ = 37° (Sill Width = 1.5 m) φ = 34° (Sill Width = 1.5 m) φ = 40° (Sill Width = 0.8 m) φ = 37° (Sill Width = 0.8 m)

74 Figure 2-61. Relationship between applied pressure and rotation of sill : integrated sill, s = 0.4 m, and medium sand foundation. -1 0 1 2 3 4 5 0 100 200 300 400 500 600 700 800 900 1000 Applied Pressure (kPa) R o t a t i o n o f S i l l ( ° ) , P o s i t i v e f o r C o u n t e r - C l o c k w i s e R o t a t i o n denotes facing failure (facing failure @ 817 kPa) (facing failure @ 667 kPa) (facing failure @ 567 kPa) (facing failure @ 917 kPa) (facing failure @ 767 kPa) (facing failure @ 650 kPa) φ = 37° (Sill Width = 1.5 m) φ = 40° (Sill Width = 1.5 m) φ = 34° (Sill Width = 2.5 m) φ = 37° (Sill Width = 2.5 m) φ = 40° (Sill Width = 2.5 m) (facing failure @ 969 kPa) φ = 37° (Sill Width = 0.8 m) φ = 40° (Sill Width = 0.8 m) φ = 34° (Sill Width = 1.5 m) φ = 34° (Sill Width = 0.8 m)

75 Figure 2-62. Relationship between applied pressure and rotation of sill: isolated sill, s = 0.2 m, and medium sand foundation. -5 -4 -3 -2 -1 0 1 2 3 0 100 200 300 400 500 600 700 800 900 1000 R o t a t i o n o f S i l l ( ° ) , P o s i t i v e f o r C o u n t e r - C l o c k w i s e R o t a t i o n φ = 34° (Sill Width = 0.8 m) φ = 37° (Sill Width = 0.8 m) φ = 40° (Sill Width = 0.8 m) φ = 34° (Sill Width = 1.5 m) φ = 37° (Sill Width = 1.5 m) φ = 40° (Sill Width = 1.5 m) φ = 34° (Sill Width = 2.5 m) φ = 37° (Sill Width = 2.5 m) φ = 40° (Sill Width = 2.5 m)

76 Figure 2-63. Relationship between applied pressure and rotation of sill: isolated sill, s = 0.4 m, and medium sand foundation. -5 -4 -3 -2 -1 0 1 2 3 0 100 200 300 400 500 600 700 800 900 1000 Applied Pressure (kPa) R o t a t i o n o f S i l l ( ° ) , P o s i t i v e f o r C o u n t e r - C l o c k w i s e R o t a t i o n φ = 34° (Sill Width = 0.8 m) (facing failure @ 937 kPa) φ = 34° (Sill Width = 1.5 m) φ = 40° (Sill Width = 0.8 m) φ = 37° (Sill Width = 0.8 m) denotes facing failure (facing failure @ 717 kPa) (facing failure @ 617 kPa) (facing failure @ 533 kPa) (facing failure @ 867 kPa) (facing failure @ 750 kPa) (facing failure @ 633 kPa) φ = 40° (Sill Width = 2.5 m) φ = 37° (Sill Width = 2.5 m) φ = 34° (Sill Width = 2.5 m) φ = 40° (Sill Width = 1.5 m) φ = 37° (Sill Width = 1.5 m)

77 0 2 4 6 8 10 12 14 16 18 0 100 200 300 400 500 600 700 800 900 1000 Applied Pressure (kPa) S i l l S e t t l e m e n t / L o w e r W a l l H e i g h t ( % ) φ = 34° (Sill Width = 0.8 m) φ = 40° (Sill Width = 2.5 m) φ = 37° (Sill Width = 2.5 m) φ = 34° (Sill Width = 2.5 m) φ = 40° (Sill Width = 1.5 m) φ = 37° (Sill Width = 1.5 m) φ = 34° (Sill Width = 1.5 m) φ = 40° (Sill Width = 0.8 m) φ = 37° (Sill Width = 0.8 m) Figure 2-64. Relationship between applied pressure and sill settlement: integrated sill, s = 0.2 m, and rigid foundation.

78 0 2 4 6 8 12 14 16 18 0 100 200 300 400 500 600 700 800 900 1000 Applied Pressure (kPa) S i l l S e t t l e m e n t / L o w e r W a l l H e i g h t ( % ) (facing failure @ 850 kPa) denotes facing failure φ = 34° (Sill Width = 0.8 m) (facing failure @ 969 kPa) φ = 37° (Sill Width = 0.8 m) φ = 34° (Sill Width = 1.5 m) φ = 40° (Sill Width = 0.8 m) (facing failure @ 667 kPa) (facing failure @ 550 kPa) (facing failure @ 933 kPa) (facing failure @ 767 kPa) (facing failure @ 633 kPa) φ = 40° (Sill Width = 2.5 m) φ = 37° (Sill Width = 2.5 m) φ = 34° (Sill Width = 2.5 m) φ = 40° (Sill Width = 1.5 m) φ = 37° (Sill Width = 1.5 m) Figure 2-65. Relationship between applied pressure and sill settlement: integrated sill, s = 0.4 m, and rigid foundation.

79 0 2 4 6 8 10 12 14 16 18 0 100 200 300 400 500 600 700 800 900 1000 Applied Pressure (kPa) S i l l S e t t l e m e n t / L o w e r W a l l H e i g h t ( % ) φ = 37° (Sill Width = 0.8 m) φ = 40° (Sill Width = 0.8 m) φ = 34° (Sill Width = 1.5 m) φ = 37° (Sill Width = 1.5 m) φ = 40° (Sill Width = 1.5 m) φ = 34° (Sill Width = 2.5 m) φ = 37° (Sill Width = 2.5 m) φ = 40° (Sill Width = 2.5 m) φ = 34° (Sill Width = 0.8 m) Figure 2-66. Relationship between applied pressure and sill settlement: isolated sill, s = 0.2 m, and rigid foundation.

80 0 2 4 6 8 10 12 14 16 18 0 100 200 300 400 500 600 700 800 900 1000 Applied Pressure (kPa) S i l l S e t t l e m e n t / L o w e r W a l l H e i g h t ( % ) (facing failure @ 700 kPa) denotes facing failure φ = 34° (Sill Width = 0.8 m) φ = 40° (Sill Width = 0.8 m) φ = 34° (Sill Width = 1.5 m) φ = 37° (Sill Width = 1.5 m) φ = 40° (Sill Width = 1.5 m) φ = 34° (Sill Width = 2.5 m) φ = 37° (Sill Width = 2.5 m) φ = 40° (Sill Width = 2.5 m) (facing failure @ 600 kPa) (facing failure @ 717 kPa) (facing failure @ 817 kPa) (facing failure @ 500 kPa) (facing failure @ 583 kPa) φ = 37° (Sill Width = 0.8 m) (facing failure @ 906 kPa) Figure 2-67. Relationship between applied pressure and sill settlement: isolated sill, s = 0.4 m, and rigid foundation.

81 0 2 4 6 8 10 12 0 100 200 300 400 500 600 700 800 900 1000 Applied Pressure (kPa) M a x i m u m L a t e r a l W a l l M o v e m e n t / L o w e r W a l l H e i g h t ( % ) φ = 37° (Sill Width = 0.8 m) φ = 40° (Sill Width = 0.8 m) φ = 34° (Sill Width = 1.5 m) φ = 37° (Sill Width = 1.5 m) φ = 40° (Sill Width = 1.5 m) φ = 34° (Sill Width = 2.5 m) φ = 37° (Sill Width = 2.5 m) φ = 40° (Sill Width = 2.5 m) φ = 34° (Sill Width = 0.8 m) Figure 2-68. Relationship between applied pressure and maximum lateral wall displacement: integrated sill, s = 0.2 m, and rigid foundation.

82 Figure 2-69. Relationship between applied pressure and maximum lateral wall displacement: integrated sill, s = 0.4 m, and rigid foundation.

83 Figure 2-70. Relationship between applied pressure and maximum lateral wall displacement: isolated sill, s = 0.2 m, and rigid foundation.

84 Figure 2-71. Relationship between applied pressure and maximum lateral wall displacement: isolated sill, s = 0.4 m, and rigid foundation.

85 Figure 2-72. Relationship between applied pressure and sill lateral movement: integrated sill, s = 0.2 m, and rigid foundation.

86 Figure 2-73. Relationship between applied pressure and sill lateral movement: integrated sill, s = 0.4 m, and rigid foundation.

87 Figure 2-74. Relationship between applied pressure and sill lateral movement: isolated sill, s = 0.2 m, and rigid foundation.

88 Figure 2-75. Relationship between applied pressure and sill lateral movement: isolated sill, s = 0.4 m, and rigid foundation.

89 Figure 2-76. Relationship between applied pressure and rotation of sill: integrated sill, s = 0.2 m, and rigid foundation. R o t a t i o n o f S i l l ( ° ) , P o s i t i v e f o r C o u n t e r - C l o c k w i s e R o t a t i o n

90 Figure 2-77. Relationship between applied pressure and rotation of sill: integrated sill, s = 0.4 m, and rigid foundation. R o t a t i o n o f S i l l ( ° ) , P o s i t i v e f o r C o u n t e r - C l o c k w i s e R o t a t i o n

91 Figure 2-78. Relationship between applied pressure and rotation of sill: isolated sill, s = 0.2 m, and rigid foundation. R o t a t i o n o f S i l l ( ° ) , P o s i t i v e f o r C o u n t e r - C l o c k w i s e R o t a t i o n

92 Figure 2-79. Relationship between applied pressure and rotation of sill: isolated sill, s = 0.4 m, and rigid foundation. -5 -4 -3 -2 -1 0 1 2 3 0 100 200 300 400 500 600 700 800 900 1000 Applied Pressure (kPa) R o t a t i o n o f S i l l ( ° ) , P o s i t i v e f o r C o u n t e r - C l o c k w i s e R o t a t i o n (facing failure @ 906 kPa) φ = 37° (Sill Width = 0.8 m) (facing failure @ 500 kPa) (facing failure @ 817 kPa) (facing failure @ 717 kPa) (facing failure @ 600 kPa) φ = 37° (Sill Width = 2.5 m) φ = 34° (Sill Width = 2.5 m) φ = 40° (Sill Width = 1.5 m) φ = 37° (Sill Width = 1.5 m) φ = 34° (Sill Width = 1.5 m) φ = 40° (Sill Width = 0.8 m) φ = 34° (Sill Width = 0.8 m) φ = 40° (Sill Width = 2.5 m) denotes facing failure (facing failure @ 700 kPa) (facing failure @ 583 kPa)

93 • The reinforcement spacing and sill width affects the lat- eral movement of the sill significantly. With an isolated sill, the lateral movement of the sill is nearly indepen- dent of soil friction angle for a small sill width (0.8 m). The effect became slightly more pronounced when sill width became larger. • A major difference between the integrated sill and iso- lated sill is in the rotation of the sill. The integrated sills all experienced counter-clockwise tilting (positive val- ues of rotation in the figures), while the isolated sills generally experienced clockwise rotation (negative val- ues of rotation in the figures), except for sill width = 2.5 m, where the rotations were clockwise. • With a rigid foundation, the abutments tended to have significantly smaller sill settlements, smaller maximum lateral wall displacements, smaller sill lateral move- ments, and smaller sill rotations (except for isolated sills) than the abutments situated over a medium sand foundation. Allowable Bearing Pressures The allowable bearing pressures of GRS abutments were evaluated using the results of the 36 analyses that were done with a medium sand foundation, because they offered more conservative allowable bearing pressures than those with a rigid foundation. Two performance criteria were exam- ined. One criterion involved a limiting sill settlement, where the allowable bearing pressure corresponded to a sill settlement of 1 percent of the lower wall height (i.e., 1 percent H). The other criterion involved distribution of the critical shear strain in the reinforced soil mass, where the allowable bearing pres- sures corresponded to a condition in which a triangular crit- ical shear strain distribution reached the back edge of the sill (i.e., heel of the sill). The 1 percent H criterion came from the existing maximum settlement criteria for bridge abutments proposed by Bozozuk (1978), Walkinshaw (1978), Grover (1978), and Wahls (1990). The existing settlement criteria for ride quality and structural distress range from 51 mm to 102 mm based on experience with real bridges. Given that the finite element analysis results should be regarded as short-term response, and the long-term settlement is likely to be of about the same magnitude as the short-term settlement (see “Assessment of the NCHRP Test Abutments” in Chap- ter 3), 1 percent H or 47 mm short-term settlement as obtained from the analysis was adopted as a criterion for evaluating the allowable bearing pressures. The critical shear strain concept has been used in the Cam clay model, a widely used soil model developed at Cam- bridge University in the United Kingdom, for assessing fail- ure of a soil mass. The critical shear strain for the three soils used in the analysis was 3.2 percent, as determined from the triaxial test results. A more detailed explanation of the criti- cal shear strain distribution criterion is in the parametric study in this chapter. Tables 2-7 and 2-8 show the values of the bearing pres- sures corresponding to the 1 percent H settlement criterion and the critical shear strain distribution criterion for all 36 analyses with a medium sand foundation. The critical shear strain distribution criterion generally yields a somewhat higher allowable bearing pressure for reinforcement spacing of 0.2 m than that with the 1 percent H settlement criterion. This observation, however, is less consistent for reinforce- ment spacing of 0.4 m. The values of bearing pressures presented in Tables 2-7 and 2-8 were used as the basis for the recommended allowable bearing pressures in the recom- mended design method (Chapter 3).

94 TABLE 2-7 Allowable bearing pressures based on the 1%H settlement criterion φ (degrees) Sill Type Sill Width (m) Reinforcement Spacing (m) Applied Pressure at Settlement = 1%H (kPa) 34 259 37 3150.2 40 354 34 231 37 284 0.8 0.4 40 324 34 162 37 1920.2 40 221 34 145 37 175 1.5 0.4 40 202 34 132 37 1540.2 40 175 34 121 37 143 Integrated 2.5 0.4 40 162 34 207 37 2420.2 40 268 34 166 37 201 0.8 0.4 40 225 34 132 37 1570.2 40 180 34 109 37 132 1.5 0.4 40 150 34 110 37 1310.2 40 148 34 91 37 115 Isolated 2.5 0.4 40 132

95 Allowable Bearing φ (degrees) Pressure (εcritical = 3.2%) Sill Type Sill Width (m) Reinforcement Spacing (m) (kPa) 34 281 37 3750.2 40 500 34 156 37 281 0.8 0.4 40 406 34 167 37 2170.2 40 283 34 117 37 167 1.5 0.4 40 233 34 150 37 2000.2 40 267 34 100 37 150 Integrated 2.5 0.4 40 217 34 188 37 2500.2 40 313 34 125 37 156 0.8 0.4 40 219 34 150 37 2000.2 40 250 34 83 37 133 1.5 0.4 40 167 34 133 37 1670.2 40 233 34 67 37 117 Isolated 2.5 0.4 40 150 TABLE 2-8 Allowable bearing pressures based on the critical shear strain distribution criterion

Next: Chapter 3 - Interpretation, Appraisal, and Applications »
Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing Get This Book
×
MyNAP members save 10% online.
Login or Register to save!
Download Free PDF

TRB's National Cooperative Highway Research Program (NCHRP) Report 556: Design and Construction Guidelines for Geosynthetic-Reinforced Soil Bridge Abutments with a Flexible Facing, presents the findings of research undertaken to develop a rational design method and construction guidelines for using geosynthetic-reinforced soil (GRS) systems in bridge abutments. The report includes two appendixes. A third appendix, "Verification of the Analytical Model, " is available as NCHRP Web-Only Document 81.

  1. ×

    Welcome to OpenBook!

    You're looking at OpenBook, NAP.edu's online reading room since 1999. Based on feedback from you, our users, we've made some improvements that make it easier than ever to read thousands of publications on our website.

    Do you want to take a quick tour of the OpenBook's features?

    No Thanks Take a Tour »
  2. ×

    Show this book's table of contents, where you can jump to any chapter by name.

    « Back Next »
  3. ×

    ...or use these buttons to go back to the previous chapter or skip to the next one.

    « Back Next »
  4. ×

    Jump up to the previous page or down to the next one. Also, you can type in a page number and press Enter to go directly to that page in the book.

    « Back Next »
  5. ×

    To search the entire text of this book, type in your search term here and press Enter.

    « Back Next »
  6. ×

    Share a link to this book page on your preferred social network or via email.

    « Back Next »
  7. ×

    View our suggested citation for this chapter.

    « Back Next »
  8. ×

    Ready to take your reading offline? Click here to buy this book in print or download it as a free PDF, if available.

    « Back Next »
Stay Connected!