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‘Old-timer’ recalls the history of geomembrane interface strength tests (Part 1)

February 1st, 2011 / By: / Feature, Geomembranes, Testing & Codes

Introduction

I was fortunate enough to attend the interesting 9th International Conference on Geosynthetics (9ICG) in Guarujá, Brazil, in May 2010.

During the last day of presentations there was a good discussion on large-scale, laboratory, direct-shear liner testing: Whether to clamp and restrain the liner sheet or allow it to “float” unrestrained. This specific give-and-take brought back many fond memories for me on this subject and as well as numerous other laboratory and field testing issues that liner engineers faced in the early days of the 1980s into the 1990s.

I mentioned this to the editor of this magazine, saying that it would be nice if some of the “old-timers” (engineers, technicians, resin suppliers, manufacturers, and installers) would start writing down some of their experiences and publish historic articles to help the “youngsters” understand how we have arrived at where we are today with liner testing, design, and construction practices.

I am now 60, and have been involved in liner design and construction on more than 500 projects since the mid-1980s on the mining side (lined heap leach pads, process ponds, tailings dams, etc.). I guess that qualifies me as one of the old-timer engineers who can provide some historic perspective and maybe some insight from my file notes on laboratory direct shear testing vs. the real world conditions.

I would like to acknowledge that I got the inspiration for writing this article from the 9IGC post-conference direct shear box discussions mentioned above (restrained vs. floating unrestrained liners) among several participants, including Ian Peggs, J.P. Giroud, and Sam Allen, among others who are well-respected in the liner community.

Background on lined slope failures

Actual slope failures on liner systems are the true tests for determining liner interface friction strengths.

Engineers can only hope to simulate planned construction and operation conditions as much as practical in the laboratory strength tests. But a slide failure allows us to back-calculate the true interface friction strength from the pre-failure and post-failure geometry, with material conditions at the time of failure. Therefore, it is worth summarizing a few of the past slope failures that have occurred on geomembrane-lined fill structures that included solid waste landfills, heap leach pads, tailings dams, and reclamation cover fill caps.

Going back to the 1970s and ’80s, geomembrane liners were then used to line waste fill and leach pad impoundments. Some of the lined impoundments, particularly for leach pads, included a sand and gravel drain cover fill placed above the liner surface for drainage or for exposure protection from climatic conditions.

The liner drain fill covers began to fail as early as 1983 on some of the steeper valley sideslopes of the impoundments during construction or in subsequent downhill high-fill lift placement operations. Dozers would slide backward to the bottom of the valley slopes while placing thin drain fill covers, or large haul trucks would slip on the lined access road ramps while bringing ore heap fill to the valley bottom for controlled thick-lift fill placement.

In an early case of a lined uranium tailings dam, constructed in 1976 with relatively flat 3.5H:1V side slopes and covered with about 2ft (0.6m) of slope drain fill cover, the water pool wave action on the lower impoundment slopes eroded the drain fill cover, causing the unsupported upper slope liner cover fill to slide downward and tear the liner.

This was worrisome enough, but by the late 1980s, several slope failures had occurred beneath geomembrane lined landfills and leach pad ore fills along relatively flat liner bottom grades. The valley bottom planar slides generally occurred at the low permeability composite soil or geotextile contact directly beneath the geomembrane liner.

Most of the engineering community first became aware of the potential for low planar interface strengths in geomembrane liner systems following the 1988 Kettleman Hills landfill slope failure in California (Mitchell et al., 1990). Several major landfill slope failures occurred between 1988 and 1997 in North America, Europe, Africa, and South America (Koerner and Soong, 1999).

Other lesser-known leach pad slope failures occurred between 1985 and 1993 at mine sites in North America, South America, and Australia (Breitenbach, 1997). However, as worrisome as this was becoming, the Northridge earthquake in Southern California in 1994 (Matasovic et al., 1995) and subsequent earthquakes in Chile and Peru have given engineers in high-seismicity countries some confidence that earthquakes have a low seismic impact on lined and fully drained fill structures.

(An interesting side note to the author’s 1997 paper was that its Abstract was rejected for the 1995 geosynthetics conference in Nashville, Tenn. After more than 30 leach pad slope failures, and some much needed arm twisting help from Rob Swan at Geosyntec, my Abstract and ensuing paper were accepted for presentation at the 1997 geosynthetics conference in Long Beach, Calif.)

The first conference I can recall with a session focused on liners in mining was the December 2005 GRI-19 conference in Las Vegas, more than 20 years after the first mining liner slope failures had occurred (leach pad liner failures, several years before the Kettleman Hills landfill failure).

I recall our discussions at the 2005 GRI conference included leach pad drain pipe failures and the effects of soil arching pipe load transfer onto the underlying liner systems, based on case histories and large-scale load tests performed more than 10 years earlier. Bob Koerner expressed his surprise about the “new” drain pipe issues on liners, and mining engineers owe him our thanks and appreciation for now seeing geosynthetics conference sessions on liners in mining, such as the 2008 GeoAmericas conference in Cancún, Mexico, and also the 9ICG in Brazil.

So fully drained fills under dynamic seismic loading were found to be a non-issue, while partially drained or saturated fills and foundations became a big static slope stability issue. The known static slope failures from excessive hydraulic conditions above the liner system included a combination of one or more of the following:

  • intense rain storm events.
  • poor internal containment material drainage.
  • excess solution application on the fill surface.
  • leachate injection into the fill material.
  • solution pipeline breaks near exterior fill slopes.

The known static slope failures from various weak foundation conditions beneath the liner system included a combination of one or more of the following:

  • re-activating an existing landslide surface.
  • thawing of frozen subgrade soils.
  • overly wet subgrade (natural soils or underliner fills).
  • subgrade subsidence in compressible natural soils and low density fills.
  • excavations at critical downgradient fill toe areas that caused unloading of toe support materials.

The design engineer needed quick and reliable test data to solve these problems, and the direct shear box test, along with other geotechnical strength tests, were under consideration for coming up with the right answers.

(Although Karl Terzaghi, Arthur Casagrande, Ralph Peck, and other geotechnical pillars in the engineering world would certainly have been scratching their heads about engineers using an impervious geomembrane liner in contact with a wet of optimum low-permeability clayey soil in a direct shear test apparatus with no way to measure pore pressure conditions!)

Background on direct shear box testing

The direct shear box was originally intended as a quick and simple test for measuring the effective shear strength of relatively free-draining noncohesive sandy soils (no excess pore pressure conditions).

The first direct shear tests were actually conducted centuries ago by Coulomb in 1776. Specific laboratory direct shear test procedures were established by 1951 (Lambe and Whitman, 1969).

The commonly accepted shear box size for noncohesive sandy soil testing was a 4in. x 4in. (100mm x 100mm) square box perfected by Ralph Peck in 1954. The 4-in.-square direct shear box quickly became the “unofficial” test method throughout the 1980s into the ’90s for determining liner interface strengths along the planar weak contact between the fine grained soils and the smooth geomembrane liner surface.

Other types of liner strength test methods were also adopted to a lesser extent, including circular ring shear tests and liner “pullout” shear tests. The consistency of the 4in. direct shear liner test results in the 1980s was erratic with no consistent laboratory test standards at the few laboratories in the U.S. capable of accurately performing the tests. The size of the box was too small for repeatable test results using the same laboratory with the same technician preparing the same sample material for testing (during my Westec days in the early 1990s with Russ Browne and Mike Henderson).

In addition, there was a significant reduction in the shear strength of the more semicrystalline (more rigid) smooth HDPE liner sheets compared to the less semicrystalline (more flexible) PVC, CSPE, and PP liners. The manufacturer’s experiments to improve on the HDPE smooth sheet low strengths led to the development of various types of higher-strength surface textured HDPE sheets by the early 1990s.

This certainly added more complexity to the testing in the small direct shear box as well as other liner sheet integrity issues. High confining loads were still being applied at zero consolidation time to all the reported test values in the geosynthetics textbooks and in technical papers at the time that I attended the 1993 geosynthetics conference in Vancouver, B.C., Canada.

(The strength values tabulated in geosynthetics textbooks for any composite liner system that involves clayey soils and pore pressure are currently of little practical use for student reference. That is, unless the basic testing support information can be provided, including, as a minimum, the shear box sizes greater than 4in., with a detailed description of the geomembrane liner type/thickness/texture, overliner and underliner contact material gradations, plasticity, moisture/density sample preparation, and loading time/shear rate conditions. Tables 1 and 2 list the basic direct shear box test information I tabulated in the early 1990s for guidance on leach pad designs because there were a significant number of pad liner slip failures occurring long before the 1993 Vancouver geosynthetics conference.) Table 1  Underliner and overliner material properties Table 2  Direct shear strength test results

From my memory, there were very few, if any, conference discussions throughout the 1990s concerning pore pressure effects and related confining stress load times prior to or during direct shear box testing. Even recently in the 9ICG conference in Brazil, I listened to a speaker essentially saying that “the geomembrane liner contact with the clayey soil acts as a more pervious zone (inferring minimal intimate liner contact related to liner folds, or that there are no composite liner designs that work?) for minimizing any direct shear pore pressure problems.”

I think it is time to re-educate the youngsters that pore pressures are present in any moderate- to high-plasticity clayey soils placed in the shear box at above optimum moisture content and instantaneously loaded. In addition, excess pore pressures are one of the primary factors causing most of the heap leach pad slope failures to occur during placement of the first ore lift load (highest change in load stress condition) and not at some higher ultimate stack load!

Real world construction also tells us to expect a wet liner contact surface to develop as liners are exposed to night and day temperature fluctuations (moisture is drawn upward to the underliner contact surface from heating and cooling moisture condensation). The temperature fluctuation on exposed liner surfaces also relates to above optimum moisture desiccation cracks occurring in the clayey underliner soils, but that is another subject for liner engineers and regulator to talk about.

The earliest time that I recall using delayed pre-test load consolidation times for studying pore pressure effects was about 1991with my best lab technician in the world, Rob Swan, at Geosyntec’s laboratory in Atlanta.

We were experimenting with the larger 6in. (150mm) and 12in. (300mm) direct shear boxes as early as 1989 for several high fill load heap leach projects in Idaho at that time. Some of the Atlanta lab testing also included gluing PVC liner to wood blocks to isolate the total stress pore pressure issue for another day and focus on determining the absolute lowest zero pore pressure planar effective stress strength on the very flexible liner surfaces under high loads (along with the complaints from Rob that the glue was not holding at the 300psi confining pressure shear loads).

He, indeed, survived the tests and I do owe him a great deal of thanks for his hard efforts. I actually had him thanking me at the time for having him do these challenging leach pad strength tests that were a technical lab challenge and with loading conditions far beyond anything he was doing for the landfill industry.

Bigger 6in. (150mm) and 12in. (300mm)-square direct shear box tests became available at several other liner testing laboratories in the early 1990s to accommodate composite and textured liner designs, which also allowed the use of larger liner cover fill angular rock sizes of typically ¾in. (19mm) sieve size up to larger 2in. (50mm) maximum rounded river rock sizes.

The 12-in.-square direct shear box test procedures for the liner interface strengths were officially established in November 1992 (ASTM D5321), thanks to the long and hard efforts of geomembrane liner direct shear testing pioneers like Greg Richardson, Sam Allen, and Rob Swan. Although the 12-in.-square shear box and established test standards produced more consistent test results compared to the 4-in.-square shear box, other factors needed consideration in applying simulated laboratory test strengths to actual site conditions during construction and operation of the lined facilities.

Direct shear testing parameters

General

Based on numerous large-scale laboratory direct shear box tests in the 1990s by Westec on high-load heap leach liners at delayed consolidation times, and running a battery of tests at varying placed clayey soil moisture contents, the geomembrane liner interface contact with underlying and overlying fill materials showed a distinct increase in friction strength vs. load consolidation time (Breitenbach and Swan, 1999).

The increase in liner interface peak and residual friction strength with respect to time was mainly due to two conditions:

  1. the apparent effect of high-load deformations or “dimpling” on a micro scale into the planar elastic geomembrane liner surface contact with overlying and underlying soil rock particles and
  2. a reduction in low permeability underliner soil excess pore pressure conditions with an increase in pre-test load consolidation time.

The measured cohesion or apparent adhesion strengths in the study decreased with time, and therefore were assumed to be negligible for conservative long-term leach pad liner strength conditions.

The larger direct shear box apparatus first used in 1989 and pre-test load consolidation time in 1991 became my primary tools in the early Westec years of liner design for determining more accurate liner strengths (helped to explain what was happening to leach pad failures in the real world).

The other big challenge was to maintain optimum or dry of optimum moisture conditions in the underliner fill for low pore pressure conditions while maintaining a low permeability requirement (this was difficult to do with low permeabilities from the laboratory fixed wall permeameter tests shown to drop an order of magnitude when wetted a few percent above optimum moisture content).

Since the first ore lift was typically placed at 5–10m (16–32ft) of load height, I used this equivalent fill loading in laboratory permeability tests to achieve an acceptable low permeability value without compromising the stability of the leach pad from wet of optimum soils subjected to rapid loading ore lift conditions. The engineering challenge that still remains today is how to accurately apply the laboratory direct shear test results to real world site conditions with the other variable testing parameters listed below.

Parameters influencing liner test strengths

Published information on geomembrane liner interface shear strengths in textbooks generally present a brief description of test sample materials (soil overliner/underliner classification, liner type, liner thickness, smooth or textured liner surface) along with the peak or residual friction angle and apparent cohesion interface strength test results in a table for quick and easy reference.

This is insufficient information for interpreting the shear test results. For the engineer to effectively use this information as a reference guideline for assumed design strengths on a particular project with similar liner and fill materials, there are more factors to consider.

As an example, the underliner soil compaction could have been prepared as either standard density (ASTM D698) or modified density (ASTM D1557) at a remolded moisture content above or below optimum moisture content with a specified dry density typically ranging from 90–100% compaction.

If the underliner soil compaction preparation procedures are not reported along with the liner test strengths, the test results at 90% of the lower standard density test method vs. 100% of the higher modified density test method could cause the published test strength values to significantly vary and render the test data useless.

Furthermore, if the lower standard density test is placed at 2% wet of optimum moisture content and the higher modified density test is placed at 2% dry of optimum moisture content, then the published test strength values could significantly vary by more than 20%.

These examples demonstrate many reasons why the published strengths are highly variable for similar soil materials and liner types. In fact, there are so many factors influencing the laboratory direct shear interface strengths to where site-specific material tests are recommended as a requirement along with appropriate laboratory test equipment and procedures for final engineering analyses and designs.

In reviewing my early 1990 notes on direct shear testing, there were several interrelated factors (and possibly many more that can be added to the list by laboratory soils technicians) that I considered to have the greatest influence on the liner interface test strengths, as listed below in no particular order of importance:

  • test equipment size (4in. vs. 12in. shear box size).
  • rate of normal confining stress loading (low vs. high loads / pre-test load time).
  • rate of applied shear force (slow vs. fast).
  • geomembrane liner thickness (40–80mil typical).
  • geomembrane interface contact (underliner/overliner soil or geotextile).
  • geomembrane liner anchorage (restricted vs. unrestricted).
  • geomembrane liner flexibility (more vs. less flexible).
  • geomembrane liner surface (smooth vs. textured and to what degree of texturing).
  • geomembrane liner texture type (sprayed, dripped, coextruded, or calendared).
  • underliner/overliner soil classification (gradation and plasticity).
  • underliner/overliner maximum rock size (particle distribution at the interface contact).
  • underliner/overliner rock particle shape (rounded vs. angular).
  • underliner/overliner moisture content (dry vs. wet of optimum moisture).
  • underliner/overliner compaction (low vs. high density).
  • strength condition (effective stress vs. total stress).
  • strength selection (peak vs. residual shear strength).

Part 2 of this article continues in the April/May issue of Geosynthetics.

Allan Breitenbach is Principal Geotechnical Engineer at Ausenco Vector, based in Denver, Colo., USA.

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